THE EFFECT OF INORGANIC FILLERS ON THE PROPERTIES OF
WOOD PLASTIC COMPOSITES
A Dissertation
Submitted to the Graduate Faculty of the
Louisiana State University and
Agricultural and Mechanical College
in partial fulfillment of the
requirements for the degree of
Doctor of Philosophy
In
The School of Renewable Natural Resources
by
Birm June Kim
B.S., Kookmin University, February 1997
M.S., Seoul National University, February 2004
May 2012
ACKNOWLEDGEMENTS
Many people contributed to the successful completion of this dissertation.
First of all, I cordially appreciate my committee chair, Dr. Qinglin Wu for his guidance,
advice, encouragement and support throughout my PhD program. His talent and diligence have
thoroughly motivated me and will shine on me in my future career.
I also greatly appreciate my committee members, Dr. Cornelis F. de Hoop, Dr. Ioan I.
Negulescu, Dr. Sun Joseph Chang, Dr. Sunggook Park, and Dr. W. David Constant. Their
suggestions and advices were invaluable and constructive in my research work and dissertation.
It has been an honor and privilege for me to have them on my committee.
I acknowledge my laboratory fellows and friends, Dr. Chengjun Zhou, Dr. Xia Guan, Mr.
Jingquan Han, Mr. Runzhou Huang, Mr. Kai Chi and Dr. Peng Li for their help on experiments.
Special thanks go to Dr. Fei Yao and Dr. Yanjun Xu, my former laboratory fellows, for their
suggestions and discussions on my research.
I am grateful to my parents, Sang-Eon Kim and Gi-Joo Kim for their love, care and
support. Also, I appreciate that my parents-in-law, Jae-Heui Park and Sook-Hee Jun for their
love, concern and encouragement.
Last but not least, I am heartily indebted to my wife, Sun-Young Park for her patience,
devotion, sacrifice, and endless love. She is the one who should get the credit for each and every
piece of success that I have achieved. She deserves it.
Financial support from USDA-NIFA Biomass Initiative Program (Award Number: 683A75-6-508) and the National Nature Science Foundation of China (Award Number:
31010103905) is greatly appreciated.
ii
TABLE OF CONTENTS
ACKNOWLEDGEMENTS ............................................................................................................ ii
LIST OF TABLES ......................................................................................................................... vi
LIST OF FIGURES ...................................................................................................................... vii
NOMENCLATURE ....................................................................................................................... x
ABSTRACT .................................................................................................................................. xii
CHAPTER 1 INTRODUCTION .................................................................................................... 1
1.1 BACKGROUND .................................................................................................................. 1
1.2 OBJECTIVES ....................................................................................................................... 3
1.3 ORGANIZATION OF DISSERTATION ............................................................................ 4
1.4 REFERENCES ..................................................................................................................... 4
CHAPTER 2 PERFORMANCE OF BAMBOO PLASTIC COMPOSITES WITH HYBRID
BAMBOO AND PRECIPITATED CALCIUM CARBONATE FILLERS ................................... 6
2.1 INTRODUCTION ................................................................................................................ 6
2.2 EXPERIMENTAL ................................................................................................................ 9
2.2.1 Materials and Preparation .................................................................................. 9
2.2.2 Composite Manufacturing................................................................................ 10
2.2.3 Material Characterizations ............................................................................... 10
2.3 RESULTS AND DISCUSSION ......................................................................................... 12
2.3.1 Basic Properties of Bamboo Filler ................................................................... 12
2.3.2 Basic Properties of PCC................................................................................... 16
2.3.3 Mechanical and Morphological Properties of Bamboo-PP/PE Composites .... 17
2.3.4 Mechanical and Morphological Properties of PCC-PP/PE Composites .......... 22
2.3.5 Properties of PCC-Bamboo-PP/PE Composites .............................................. 23
2.4 CONCLUSIONS ................................................................................................................ 32
2.5 REFERENCES ................................................................................................................... 32
CHAPTER 3 MECHANICAL AND PHYSICAL PROPERTIES OF CORE-SHELL
STRUCTURED WOOD PLASTIC COMPOSITES: EFFECT OF SHELLS WITH HYBRID
MINERAL AND WOOD FILLERS ............................................................................................ 36
3.1 INTRODUCTION .............................................................................................................. 36
3.2 EXPERIMENTAL .............................................................................................................. 39
3.2.1 Materials and Preparation ................................................................................ 39
3.2.2 Coextruded WPC Manufacturing..................................................................... 40
3.2.3 Characterization ............................................................................................... 41
3.3 RESULTS AND DISCUSSION ......................................................................................... 42
iii
3.3.1 Properties of Treated PCC ............................................................................... 42
3.3.2 Mechanical and Morphological Properties of Composites .............................. 47
3.3.3 Impact Fracture Type Comparison .................................................................. 53
3.3.4 WA and TS Properties of Composites ............................................................. 57
3.3.5 Thermal Expansion Properties of Composites ................................................. 59
3.4 CONCLUSIONS ................................................................................................................ 59
3.5 REFERENCES ................................................................................................................... 60
CHAPTER 4 MECHANICAL PROPERTIES OF CO-EXTRUDED WOOD PLASTIC
COMPOSITES WITH GLASS FIBER FILLED SHELL ............................................................ 62
4.1 INTRODUCTION .............................................................................................................. 62
4.2 EXPERIMENTAL .............................................................................................................. 64
4.2.1 Materials and Preparation ................................................................................ 64
4.2.2 Injection Molding of GF-HDPE Samples ........................................................ 65
4.2.3 Co-extruded WPC Manufacturing ................................................................... 66
4.2.4 Mechanical Testing and Characterization ........................................................ 66
4.3 RESULTS AND DISCUSSION ......................................................................................... 67
4.3.1 Shell GF-HDPE Composite ............................................................................. 67
4.3.1.1 Morphology............................................................................................... 67
4.3.1.2 Mechanical Property ................................................................................. 67
4.3.2 Extruded Core-only Composites ...................................................................... 71
4.3.2.1 Morphology............................................................................................... 71
4.3.2.2 Mechanical Property ................................................................................. 71
4.3.3 Coextruded Core-Shell Composites................................................................. 74
4.3.3.1 Morphology............................................................................................... 74
4.3.3.2 Mechanical Property ................................................................................. 77
4.3.4 Analysis of Composite Property ...................................................................... 80
4.4 CONCLUSIONS ................................................................................................................ 84
4.5 REFERENCES ................................................................................................................... 85
CHAPTER 5 SOUND TRANSMISSION PROPERTIES OF MINERAL FILLED HIGH
DENSITY POLYETHYLENE (HDPE) AND WOOD-HDPE COMPOSITES .......................... 88
5.1 INTRODUCTION .............................................................................................................. 88
5.2 THEORETICAL APPROACH .......................................................................................... 90
5.2.1 Impedance Tube ............................................................................................... 90
5.2.2 Mass Law ......................................................................................................... 95
5.2.3 Stiffness Law ................................................................................................... 96
5.3 EXPERIMENTAL .............................................................................................................. 97
5.3.1 Materials and Composite Sample Preparation ................................................. 97
5.3.2 Mechanical Property Characterization ............................................................. 99
5.3.3 Acoustic Property Characterization ................................................................. 99
5.4 RESULTS AND DISCUSSION ....................................................................................... 100
5.4.1 Basic Properties of Mineral Filled HDPE and WPC ..................................... 100
5.4.2 General TL Curves of Filled HDPE and WPCs ............................................. 101
iv
5.4.3 Comparison of TL with Mass and Stiffness Law Predictions ....................... 106
5.4.4 Effect of Filler Loading Level on Mean TL .................................................. 110
5.5 CONCLUSIONS .............................................................................................................. 112
5.6 REFERENCES ................................................................................................................. 112
CHAPTER 6 CONCLUSIONS AND FUTURE WORK ........................................................... 115
6.1 OVERALL CONCLUSIONS ........................................................................................... 115
6.2 FUTURE WORK.............................................................................................................. 117
APPENDIX: PERMISSION LETTER ....................................................................................... 119
VITA ........................................................................................................................................... 125
v
LIST OF TABLES
Table 2.1 Chemical compositions estimated from deconvoluted DTG curves for both GBP and
RBF ............................................................................................................................................... 13
Table 2.2 Mechanical properties of PCC-bamboo-PP/PE composites ......................................... 25
Table 2.3 WA (%) of variously treated bamboo-PP/PE composites with and without PCC........ 31
Table 2.4 Two-way ANOVA tests of the effect of fiber and PCC on WA of R-PP/PE composites
at different time intervals. ............................................................................................................. 31
Table 3.1 Element composition, oxygen-carbon, and silicon-oxygen ratios of untreated PCC and
TPCC from the XPS meaturement. ............................................................................................... 44
Table 3.2 Mechanical and thermal expansion properties of coextruded core-shell WPCs with
various levels of TPCC and WF in the shell layer. ....................................................................... 48
Table 4.1 Mechanical properties of core-only WPCs with different core quality (weak, moderate,
and strong) and shell-only composites with various GF content. ................................................. 69
Table 4.2 Mechanical properties of coextruded composites with different core quality (weak,
moderate, and strong) and shell thickness including various GF content in shell layer. .............. 78
Table 4.3 Two-way ANOVA tests of the effect of GF content in shell and shell thickness on
mechanical properties of different core coextruded composites................................................... 84
Table 5.1 Materials and formulations of each composite sample for acoustic testing. ................ 98
Table 5.2 Mechanical properties (MOE, MOR, and impact strength), stiffness, and area density
of mineral filled composites and WPCs as a function of filler content. ..................................... 100
Table 5.3 Summary of sound transmission property of filled HDPE and WPCs. ...................... 110
vi
LIST OF FIGURES
Figure 2.1 Thermogravimetric data of bamboo filler. (a) and (b): TG and DTG curves of raw
RBF and GBP; (c) and (d): DTG comparison between RBF and GBP before and after silane
treatment. ...................................................................................................................................... 13
Figure 2.2 FT-IR spectra (a) of RBF and GBP before and after silane treatment and schematic
diagram (b) of bamboo fiber surfaces chemically modified by silane treatment. ........................ 15
Figure 2.3 Raw PCC morphology and particle size data. (a): SEM micrograph of PCC particles;
(b): measured particle size distribution based on the laser diffraction technique. ........................ 18
Figure 2.4 Measured XRD pattern (a) and thermal decomposition data (b) of the PCC filler. .... 19
Figure 2.5 Comparison of measured flexural strength (a) and modulus (b) for R-PP/PE resin and
bamboo-PCC-PP/PE composites. ................................................................................................. 20
Figure 2.6 SEM micrographs of the impact fractured surfaces of TGBP (a) and TRBF (b)
composites..................................................................................................................................... 21
Figure 2.7 Comparison of flexural (a), tensile (b) and impact (c) properties of PCC-filled
composites as a function of PCC content. .................................................................................... 26
Figure 2.8 SEM micrographs of the impact fractured surfaces of R-PP/PE and PCC blends. (a):
R-PP/PE; (b): R-PP/PE and PCC (7 wt%); (c): R-PP/PE and PCC (30 wt%).............................. 27
Figure 2.9 SEM micrographs of the impact fractured surfaces of PCC-GBP-PP/PE composites:
(a) 6 wt% PCC and (b) 18 wt% PCC. ........................................................................................... 28
Figure 2.10 Comparison of flexural modulus (a) and tensile modulus (b) of PCC-GBP and PCCRBF composites before and after silane treatment as a function of PCC content. ....................... 29
Figure 3.1 XPS spectra of untreated PCC (a) and treated PCC (b). ............................................. 43
Figure 3.2 FT-IR spectra of PCC and TPCC (a: whole spectra and b: selected region). ............. 45
Figure 3.3 Schematic diagram of PCC surface chemically modified by silane treatment. .......... 46
Figure 3.4 Flexural (a: weak core, b: strong core) and impact (c) properties of coextruded coreshell structured WPCs-effect of TPCC content in a 15% WF-filled shell. ................................... 49
Figure 3.5 Flexural (a: weak core, b: strong core) and impact (c) properties of coextruded coreshell structured WPCs-effect of WF content in a 12% TPCC-filled shell. ................................... 50
vii
Figure 3.6 SEM micrographs of impact fractured surfaces of coextruded core-shell structured
WPC: (a) weak core with TPCC-only filled shell layer; (b) weak core with TPCC and WF filled
shell layer. ..................................................................................................................................... 52
Figure 3.7 Fracture type distributions (hinge versus complete) from impact tests of coextruded
core-shell structured WPCs. (a) and (b): weak core; (c) and (d): strong core. ............................. 55
Figure 3.8 Schematic diagram showing impact fracture process of coextruded core-shell
structured WPCs and photographs of typical impact fractured samples.(a) weak core system; (b)
strong core systm .......................................................................................................................... 56
Figure 3.9 TS and WA properties of coextruded core-shell WPCs. (a): weak core and (b) strong
core- effect of TPCC loading in a 15% WF-filled shell. (c): weak core and (d) strong core- effect
of WF loading in a 12% TPCC-filled shell................................................................................... 58
Figure 4.1 SEM micrographs of impact fractured surfaces of 30% GF filled HDPE
composites..................................................................................................................................... 69
Figure 4.2 Flexural properties (a) and impact strength (b) of GF filled shell-only composites. .. 70
Figure 4.3 SEM micrographs of core-only WPCs: weak core (a and b), moderate core (c and d)
and strong core (e and f). .............................................................................................................. 72
Figure 4.4 Flexural properties (a) and impact strength with percentage of breaking types (b) for
extruded core-only WPCs (weak, moderate, and strong core). .................................................... 73
Figure 4.5 SEM micrographs for cross section (a and b) and top view (c and d) of strong-core
coextruded WPC with 40% GF shell. ........................................................................................... 75
Figure 4.6 SEM micrographs of coextruded WPCs: moderate core with PE shell (a) and
with 40% GF shell (b); strong core with PE shell (c) and with 40% GF shell (d). ...................... 76
Figure 4.7 Comparison of MOE for coextruded composites with all core- and all shellcomposites: (a) weak core, (b) moderate core, and (c) strong core. ............................................. 81
Figure 4.8 Comparison of MOR for coextruded composites with all core- and all shellcomposites: (a) weak core, (b) moderate core, and (c) strong core. ............................................. 82
Figure 4.9 Comparison of impact strengths for coextruded composites with all core- and all
shell-composites: (a) weak core, (b) moderate core, and (c) strong core. .................................... 83
Figure 5.1 Schematic diagram of hardware setup (top) and the TL measurement in an impedance
tube (bottom)................................................................................................................................. 91
Figure 5.2 Comparisons of TL curves between clay-filled HDPE and WPCs (a) and between
PCC filled HDPE and WPCs (b). ............................................................................................... 102
viii
Figure 5.3 Comparisons of experimental TL curves of filled HDPE in comparison with mass and
stiffness law TL predictions: (a) HDPE, (b) HDPE/Clay4, (c) HDPE/Clay8, (d) HDPE/PCC20,
and (e) HDPE/PCC40. ................................................................................................................ 104
Figure 5.4 Comparisons of experimental TL curves of filled WPCs in comparison with mass and
stiffness law TL predictions: (a) WPC, (b) WPC/Clay4, (c) WPC/Clay8, (d) WPC/PCC20, and
(e) WPC/PCC40. ......................................................................................................................... 105
Figure 5.5 Comparisons of experimental TL curves of filled HDPE in comparison with combined
TL predictions: (a) HDPE, (b) HDPE/Clay4, (c) HDPE/Clay8, (d) HDPE/PCC20, and (e)
HDPE/PCC40. ............................................................................................................................ 108
Figure 5.6 Comparisons of experimental TL curves of filled WPCs in comparison with combined
TL predictions: (a) WPC, (b) WPC/Clay4, (c) WPC/Clay8, (d) WPC/PCC20, and (e)
WPC/PCC40. .............................................................................................................................. 109
Figure 5.7 Comparisons of Averaged TL values between clay filled HDPE and WPCs (a) and
PCC filled HDPE and WPCs (b) in S-region and M-region. ...................................................... 111
ix
NOMENCLATURE
ANOVA: analysis of variance
A-TL: averaged sound transmission loss
CC: calcium carbonate
CTE: coefficient of thermal expansion
DSC: differential scanning calorimetry
DTG: derivative thermal gravimetric
FTIR: fourier transform infrared
GBP: ground bamboo particle
GF: glass fiber
HDPE: high density polyethylene
IM: impact strength
LDPE: low density polyethylene
MA: maleic anhydride
MAPE: maleic anhydride grafted polyethylene
MOE: modulus of elasticity
MOR: modulus of rupture
M-region: mass controlled region
NCC: natural calcium carbonate
NF: natural fiber
NFPC: natural fiber/polymer composite
PCC: precipitated calcium carbonate
PP: polypropylene
x
PVC: poly (vinyl chloride)
RBF: refined bamboo fiber
RFE: experimental resonance frequency
RFS1: stiffness-1 resonance frequency
RFS2: stiffness-2 resonance frequency
RS1-M-IP: intersection point from stiffness-1 to mass law predictions
RS2-M-IP: intersection point from stiffness-2 to mass law predictions
R-PP/PE: recycled PP/PE
SEM: scanning electron microscope
S-region: stiffness controlled region
TG: thermogravimetric
TGBP: silane treated GBP
TL: sound transmission loss
TRBF: silane treated RBF
TS: thickness swelling
V-HDPE: virgin HDPE
WA: water absorption
WF: wood fiber or wood flour
WPC: wood plastic composite
XPS: X-ray photoelectron spectroscopy
XRD: X-ray diffraction
xi
ABSTRACT
The effect of inorganic fillers including precipitated calcium carbonate (PCC), glass fiber
(GF), and nano-clay on properties of structured WPCs was investigated.
In PCC-bamboo-polymer hybrid composites, tensile and flexural moduli were improved
with increasing PCC content. After silane treatment of bamboo, RBF-filled hybrid composites
showed better mechanical properties compared to those of GBP-filled hybrid composites. The
hybrid composites showed 3-4 times higher modulus than those of PCC-filled composites at high
PCC levels.
Various property differences were observed between weak- and strong-core coextruded
systems with shell composition changes. While the weak-core systems showed improved flexural
strengths compared to their core-only control, the strong-core systems had lowered flexural
strengths. In both systems, impact strengths increased at low shell filling levels but decreased at
high shell filling levels. Impact fracture types varied with core quality and shell filling
composition. Coextruded composites with treated PCC-filled shell showed better water
absorption (WA) property compared to core-only controls and coextruded composites with high
WF-filled shell. Plastic-only shell increased overall coefficient of thermal expansion (CTE) of
coextruded composites, but filled shells led to the CTE decreases of coextruded composites. GF
in shell behaved as an effective reinforcement for coextruded composites. The comparisons of
flexural property among different core systems show that GF reinforcements were optimized at
high GF loadings in a shell layer and GF alignments in the shell layer also played an important
role. In coextruded composites with different shell thicknesses, the flexural property enhanced
with the increase of shell bending modulus and strength at a given shell thickness. When the
flexural property of shell was less than that of core, the increase of shell thickness led to reduced
xii
flexural property. On the other hand, when the flexural property of shell was higher than that of
core, the opposite was true.
In sound transmission loss (TL) testing, the stiffness and surface density were major
factors influencing the sound insulation property of materials. The experimental TL results
showed that the addition of clay or PCC and/or wood fiber (WF) fillers led to the increases of
general resonance frequencies and TL in filled composites. However, at high filling levels,
composite stiffness decreases led to TL reduction. The experimental TL curves of filled HDPE
and WPCs were well approximated with the combined TL predictions from their corresponding
stiffness-1 and stiffness-2 TL for S-region and mass law TL for M-region.
xiii
CHAPTER 1 INTRODUCTION
1.1 BACKGROUND
Wood/natural fiber plastic composites (WPCs) have attracted considerable attention in
recent years. Many emerging applications for WPCs include decking, building, automobile, and
infrastructure (Klyosov 2007; Selke and Wichman 2004). Polyethylene (PE), polyvinylchloride
(PVC) and polypropylene (PP) hold a major share of resins used in WPC. Compared with
traditional glass fiber and mineral fillers, wood fillers are known to be less expensive, lighter,
sustainable, and less abrasive to processing machines. To enhance product performance,
significant effort has been made to study WPC’s properties as affected by raw material
compositions, wood-plastic interfacial bonding, and composite processing parameters (Lu, Wu,
and McNabb 2000; Mohanty, Misra, and Drzal 2001).
WPC is the combination form of raw materials such as fillers (organic and inorganic),
plastic, and additives (Nwabunma and Kyu 2008). Among the raw materials, fillers are solid
materials added to plastics mainly to reduce cost and improve properties (Kroschwitz 1990;
Whelan 1994). Typically, fillers are inexpensive, thus make the filled plastic composites less
expensive. Also, the fillers can lead to improved mechanical properties of filled composites by
restricting molecular mobility of plastic matrices (Fitzgerald, Ferrar, and Binga 1998) through
morphology changes (Okamoto et al. 1996). Therefore, the proper selection and use of filler
combinations are very important in balancing the cost and performance of WPC.
In general, fillers are divided into organic and inorganic fillers. Organic fillers include wood,
jute, and agriculture fibers. Inorganic fillers include minerals such as calcium carbonate (CC),
talc, and mica and synthetic fibers like glass, ceramic and carbon fibers. The effect of inorganic
fillers on the composite property strongly depends on the size, shape, content, surface
1
characteristics, aspect ratio, and dispersion of the fillers (Chan et al. 2002). Usually, the addition
of inorganic fillers into polymeric materials enhances the mechanical properties (mainly,
stiffness) and dimension stability of filled composites (Nwabunma and Kyu 2008; Wypych
2010). CC, talc, and mica are quite common inorganic fillers used in plastic industry and replace
the much more expensive plastic. Among the CCs, precipitated calcium carbonate (PCC) has a
more even and smaller size. PCC filled PP showed more significant effects on the brittleness,
ductility and toughness than talc and mica (Bramuzzo, Savadori, and Bacci 1985). PCC of sugar
origin is considered as waste and therefore, its new value-added applications in WPC are needed.
Glass fibers (GFs) and various nano-scale fillers (e.g., montmorillonite, graphite nanoflake,
carbon nanotube, etc) have been reported to enhance strength and barrier property of filled
plastic composites (Bramuzzo, Savadori, and Bacci 1985; Gong et al. 2004; Jiang et al. 2003).
For WPC applications, the use of high strength fillers is often cost-prohibitive. However, the
application of these materials at critical locations for structured WPC can lead to significant
property enhancement while maintaining competitive costs.
Coextruded wood plastic composite with a core-shell structure has been recently
developed and used to enhance performance characteristics of WPC (Jin and Matuana 2010;
Stark and Matuana 2007). By proper combination of constituting layers, one can achieve a
balance of such properties as light weight, high strength, high stiffness, wear resistance,
biological resistance, unusual thermal expansion characteristics, appearance, etc (Jin and
Matuana 2010; Kim and Wu 2012; Stark and Matuana 2007; Yao and Wu 2010). Fundamental
understanding of the interactions between shell and core layer with different structure (e.g.,
thickness) and material combinations is, however, needed to achieve desired product
performance. The incorporation of fillers in a shell layer of coextruded WPCs has positively
2
affected the mechanical properties of core-shell structured composites. However, cost and
performance effective shells are still needed to achieve desired composite performance.
Various sound barrier materials including concrete, brick, metal, plastic, wood and
composites have been used to reduce noise levels. Among the materials, viscoelastic polymer
materials show great potential for damping sound and vibration. However, most polymer
materials have lower elastic modulus and surface density, leading to poor sound insulating
performances when solely used as a sound barrier (Wang et al. 2011). WPCs, as new generation
green composites, offer advantages in relatively light weight, excellent recyclability, low toxicity
and high thermal stability. Thus, their application as sound barrier uses can offer a competitive
alternative to the conventional sound barriers. For this purpose, acoustic properties of structured
WPCs, influenced by composite formulations need to be fully understood.
1.2 OBJECTIVES
The objectives of the research described in this work are:
1) To investigate combined influence of PCC content, bamboo fiber type, and surface
silane treatment on the properties of filled bamboo plastic composites;
2) To elucidate the effect of treated PCC and WF loadings in the shell layer on the
mechanical, WA and CTE properties of coextruded WPCs with two different core systems;
3) To evaluate the effect of various GF contents in a shell layer and shell thickness
changes on the mechanical properties of coextruded WPCs in combination with three core
systems (low, moderate, and strong); and
4) To develop an experimental procedure for studying sound insulation properties of
mineral-filled solid plastic composites and WPCs using an impedance tube method, and to
3
investigate the effect of mineral type and loading levels on transmission loss (TL) properties of
the composites.
1.3 ORGANIZATION OF DISSERTATION
Chapter 1 provides an overall introduction of this research and the structure of this
dissertation.
Chapter 2 investigates the mechanical and WA properties of recycled-PP/PE composites
reinforced with different PCC content, bamboo fiber type and fiber surface treatment.
Chapter 3 studies the effect of treated PCC and WF loadings in the shell layer on the
mechanical, WA and CTE properties of coextruded WPC with two different core systems.
Chapter 4 discusses the effect of GF contents in a shell layer and shell thickness changes
on the flexural and impact properties of coextruded WPCs in combination with three core
systems (low, moderate, and strong).
Chapter 5 presents sound insulation property of HDPE composites and WPCs filled with
different mineral types and loading levels.
Chapter 6 provides overall conclusions of this dissertation.
1.4 REFERENCES
Bramuzzo, M., A. Savadori, and D. Bacci. 1985. Polypropylene Composites - FractureMechanics Analysis of Impact Strength. Polymer Composites 6: 1-8.
Chan, C.M., J.S. Wu, J.X. Li, and Y.K. Cheung. 2002. Polypropylene/calcium carbonate
nanocomposites. Polymer 43: 2981-92.
Fitzgerald, J.J., W.T. Ferrar, and T.D. Binga. 1998. Fatigue-resistant silicone elastomer
formulations. Journal of Applied Polymer Science 70: 1633-41.
Gong, F.L., M. Feng, C.G. Zhao, S.M. Zhang, and M.S. Yang. 2004. Thermal properties of
poly(vinyl chloride)/montmorillonite nanocomposites. Polymer Degradation and Stability 84:
289-94.
4
Jiang, H.H., D.P. Kamdem, B. Bezubic, and P. Ruede. 2003. Mechanical properties of poly(vinyl
chloride)/wood flour/glass fiber hybrid composites. Journal of Vinyl & Additive Technology 9:
138-45.
Jin, S., and L.M. Matuana. 2010. Wood/plastic composites co-extruded with multi-walled carbon
nanotube-filled rigid poly(vinyl chloride) cap layer. Polymer International 59: 648-57.
Kim, B.-J., and Q.L. Wu. 2012. Mechanical and Physical Properties of Core-Shell Structured
Wood Plastic Composites: Effect of Shells with Hybrid Mineral and Wood Fillers. Composite
Part B, In review.
Klyosov, A.A. 2007. Wood-plastic composites. Hoboken, New Jersey: John Wiley & Sons Inc.
Kroschwitz, J.I. 1990. Concise Encyclopedia of Polymer Science and Engineering, 1st ed:
Wiley-Interscience.
Lu, J.Z., Q.L. Wu, and H.S. Mcnabb. 2000. Chemical coupling in wood fiber and polymer
composites: A review of coupling agents and treatments. Wood and Fiber Science 32: 88-104.
Mohanty, A.K., M. Misra, and L.T. Drzal. 2001. Surface modification of natural fibers and
performance of the resulting biocomposites: An overview. Comp Interf. 8: 313-43.
Nwabunma, D., and T. Kyu. 2008. Polyolefin composites: John Wiley & Sons Inc.
Okamoto, M., Y. Shinoda, T. Okuyama, A. Yamaguchi, and T. Sekura. 1996. Synthesis and
crystallization behaviour of an intercalated compound of poly(ethylene terephthalate) and mica.
Journal of Materials Science Letters 15: 1178-9.
Selke, S.E., and I. Wichman. 2004. Wood fiber/polyolefin composites. Composites Part aApplied Science and Manufacturing 35: 321-6.
Stark, N.M., and L.M. Matuana. 2007. Coating WPCs using co-extrusion to improve durability.
In Conference for Coating Wood and Wood Composites: Designing for Durability, 1-12. Seattle,
WA,.
Wang, X., F. You, F.S. Zhang, J. Li, and S.Y. Guo. 2011. Experimental and Theoretic Studies on
Sound Transmission Loss of Laminated Mica-Filled Poly(vinyl chloride) Composites. Journal of
Applied Polymer Science 122: 1427-33.
Whelan, T. 1994. Polymer Technology Dictionary. London, UK: Chapman & Hall.
Wypych, G. 2010. Handbook of fillers, 3th ed: ChemTec Publishing.
Yao, F., and Q.L. Wu. 2010. Coextruded Polyethylene and Wood-Flour Composite: Effect of
Shell Thickness, Wood Loading, and Core Quality. Journal of Applied Polymer Science 118:
3594-601.
5
CHAPTER 2 PERFORMANCE OF BAMBOO PLASTIC COMPOSITES WITH
HYBRID BAMBOO AND PRECIPITATED CALCIUM CARBONATE FILLERS1
2.1 INTRODUCTION
Natural fibers have been widely used as reinforcing fillers in plastic composites due to
low cost, light weight and recyclability.(Hristov, Lach, and Grellmann 2004; Mohanty, Misra,
and Drzal 2001) Bamboo fiber is one of the traditionally used natural fibers due to its fast growth,
abundant availability and good mechanical properties from longitudinally aligned fiber
structure.(Han et al. 2008; Shin and Yipp 1989; Takaya and Tadashi 1970) It is considered as
one of the alternatives for wood resources and the use of bamboo fiber in composites can help
save wood resources.(Han et al. 2008) However, bamboo fibers are more brittle compared to
other natural fibers.(Okubo, Fujii, and Yamamoto 2004) Therefore, to enhance their reinforcing
effect in plastic composite, certain treatments of original bamboo fiber are needed. Alkali
(Cantero et al. 2003; Gassan and Bledzki 1999) and acid (Winandy, Stark, and Horn 2008) based
processes comprised most of current fiber pre-treatments, but the use of chemical agents and
high energy consumption of the process can lead to significant environmental problems and high
costs. Heat treatment with or without refining has also been introduced as natural fiber treatment
options. For instance, it was shown that high temperature steam treatment improved dimensional
stability of wood products through chemical modification of wood components.(Giebeler 1983;
Inoue and Norimoto 1991) Thermomechanical pulping through steam aided refining process
showed lower energy cost and increased fiber strength.(Vena 2005) While pre-treatments show
improved properties of natural fibers, the pre-treated fibers still have hydrophilic characteristics
different from hydrophobic plastics. Thus, to increase the compatibility between fiber and plastic
1
Reprint in part with permission from Polymer Composites
Kim, B.J., F. Yao, G. Han, Q. Wu. 2012. Performance of bamboo plastic composite with hybrid bamboo and
precipitated calcium carbonate fillers. Polymer Composites 33: 68-78.
6
matrix, various coupling agents have been used.(Lu, Wu, and Mcnabb 2000) Maleic anhydride
(MA) grafted polymers are most widely utilized coupling agents. These MA copolymers lead to
the compatibility increase of treated fibers with plastic matrix through ester linkages and
hydrogen bonds between fiber and anhydride groups,(Felix and Gatenholm 1991; Lu, Negulescu,
and Wu 2005) readily improved mechanical properties. The use of silane cross-linking agents
has also been tried in the composites. Nachtigall et al.(Nachtigall, Cerveira, and Rosa 2007)
tested the properties of polypropylene (PP) / wood fiber (WF) composites with same
concentration of silane and MA coupling agent. The result suggested that composites with silane
treatment showed better tensile strength, lower water absorption and more homogeneous
morphology than the composites made with MA. Kim et al.(Kim et al. 2011) further investigated
the effect of various silane coupling agents for PP with NaOH pre-treated WF and the whole
silane treated WF filled composites showed significantly improved tensile strength, flexural
strength and water absorption property compared to untreated WF filled composite.
Calcium carbonate (CC) is one of the most abundant and stable minerals, generally
divided into natural calcium carbonate (NCC) and precipitated calcium carbonate (PCC). While
NCC is acquired from quarrying in mines, PCC is manufactured by chemical processes.(Katz
and Milewski 1987) In plastic manufacturing, the use of PCC as a filler or an extender is
prevalent (Katz and Milewski 1987). The size, shape and content of PCC can affect mechanical
properties of composites by modifying micro morphology of base plastics.(Bramuzzo, Savadori,
and Bacci 1985; da Silva et al. 2002; Maiti and Mahapatro 1991; Misra, Deshmane, and Yuan
2007) Bramuzzo et al.(Bramuzzo, Savadori, and Bacci 1985) tested mechanical properties of PP
composites with several inorganic fillers and reported that PCC showed more significant effect
on brittleness, ductility and toughness of PP than other fillers. Maiti et al.(Maiti and Mahapatro
7
1991) investigated tensile properties of i-PP/PCC composite and showed modulus increase and
strength decrease with the increase of PCC loading. da Silva et al.(da Silva et al. 2002) reported
that when a small amount of PCC was added to the PP matrix, an increase in impact strength was
observed. But, at high PCC loadings, the impact strength was decreased. They mentioned that
this result was due to the difference of PCC dispersion in plastic matrix. Deshmane et al.(Misra,
Deshmane, and Yuan 2007) evaluated the difference in impact behavior from neat HDPE to CCHDPE composites and identified the reinforcement of composites through observing changes of
structure and mechanical behaviors processed under similar conditions. In their investigations,
they reported that the nucleating effect of CC reduced the spherulite size of base polymer, which
led to some deleterious effects on yield stress.
Several commercial wood plastic composite (WPC) deck boards contain a certain amount
of CC (e.g., Geo DeckTM by LDI Composites) and talc (e.g., Timber TechTM by Timber Tech
Company).(Klyosov 2007) The use of mineral fillers in the product can help replace much more
expensive plastic, increase stiffness of the filled products, and render the plastic more flame
resistant. For example, it was shown that 12-32% (w/w) filling ranges of CC with a 7-µm median
particle size increased the flexural modulus of HDPE-based WPC up to 64%. Also, the flexural
strength of 200-mesh CC (27.5%, w/w) and 325-mesh CC (27.5%, w/w) filled PP composite
with 40-mesh wood (27.5%, w/w) showed an increased trend from 38.6-44.8 to 45.5-51.5 MPa,
respectively.(Klyosov 2007) However, the use of PCC in WPC has not been reported (e.g., PCC
from sugar making process).(Echeverria and Holst 2005)
In this research, PCC of sugar origin and bamboo fiber were used as fillers for natural
fiber plastic composites from recycled polypropylene and polyethylene (R-PP/PE) resin. The
8
objectives of the research were to investigate combined influence of PCC content, bamboo fiber
type and silane treatment on the properties of PCC filled bamboo plastic composites.
2.2 EXPERIMENTAL
2.2.1 Materials and Preparation
R-PP/PE resin, ground bamboo particle (GBP, ≤ 20 mesh), thermo-mechanically refined
bamboo fiber (RBF, ≤ 20 mesh) and PCC were used as raw materials. Originally, R-PP/PE
(about 95% PP and 5% PE) was a commingled plastic in a fluffy form and it was pelletized using
a twin-screw extrusion machine before being blended with other raw materials. GBP was
prepared from bamboo flakes through grinding and screening using a granulator with a 20 mesh
screen. RBF was produced by refining bamboo chips at 160°C for 10 min using a KRK-2503
steam-aided disc refining system (Kumagai Riki Kogyo Co., Tokyo, Japan). Silane coupling
agent (Z-6094) from the DOW Corning Co. (Midland, MI, USA) was utilized to treat GBP and
RBF. The composition of silane was aminoethyl-aminopropyl-trimethoxysilne (> 60 wt%) and
the density of this organo-silane was 1.02 g/cm3.
Silane treatments of GBP and RBF were made by immersion methods. GBP and RBF
were firstly oven-dried at 85°C for 24 hr to reach moisture content level less than 2%. They were
immersed in an aqueous solution (methanol:water = 9:1) including silane coupling agent (3 % of
the weight of bamboo fiber) at 25°C for 1 hr. Before treatments, the silane aqueous solution was
pre-hydrolyzed for over 30min. After immersion, silane treated bamboo fiber was dried at 85°C
for 24 hr again. The treated fiber was stored in plastic bags prior to uses. PCC was obtained from
Domino Sugar Company (New Orleans, LA, USA). It was dried at 85°C for 24 hr and screened
with a 100-mesh screen.
9
2.2.2 Composite Manufacturing
A CW Brabender Intelli-Torque Twin-Screw Extruder (CW Brabender Instruments,
South Hackensack, NJ) was used to blend R-PP/PE, bamboo fiber and PCC hybrid composites.
R-PP/PE and PCC were blended first as a base material (50 wt% of plastic and 50 wt% of
PCC). After that, the pellets were diluted in the variation of PCC (6, 12, 18 wt%, respectively)
and re-compounded with bamboo fiber (40 wt%) in the blends of R-PP/PE and PCC (60 wt%)
and then extruded in the second step. The blending temperature profile was 155°C, 175°C,
180°C, 180°C, and 170°C from the feeding zone to die and the extruder rotation speed was 90
rpm. The extruded blends were pelletized and then dried in an oven at 85°C for 24 hr.
Standard test samples were made through a Battenfeld 35-Ton Plus Injection Molding
Machine (Wittmann Battenfeld GmbH, Kottingbrunn, Austria) at an injection temperature of
175°C.
2.2.3 Material Characterizations
Thermal properties of bamboo fiber and PCC were measured using TA Q-50 thermo
gravimetric analyzer (TGA: TA Instruments, New Castle, DE). Samples were heated from 25°C
to 600°C for bamboo fibers and from 25°C to 875°C for PCC at a heating rate of 10°C /min
under nitrogen flow.
FT-IR analyses of bamboo fibers were performed using a Nicolet 650 analyzer (Nicolet
Instrument Corporation, Madison, WI) to investigate the changes in the compositions of GBP
and RBF before and after treatments. Specimens were prepared as KBr pellets (bamboo fiber:
KBr = 1 : 10) and then analyzed in the range of 525–4000 cm-1. The acquired FT-IR data were
from the averages of 64 scans.
10
The morphologies of PCC and composites were analyzed by a Hitachi S-3600N VP
Scanning Electron Microscope (SEM) (Hitachi Ltd., Tokyo, Japan). The samples were coated
with Pt to improve the surface conductivity before observation and observed at an acceleration
voltage of 15 kV.
The surface area of PCC was acquired from the multipoint Brunauer, Emmett and Teller
(BET) method based on the isothermal adsorption of nitrogen. An Autosorb-1 Surface Area and
Pore Size Analyzer (Quantachrome Instruments, Boynton Beach, FL) was used. Dried PCC
particles were degassed first and then measured using nitrogen gas.
The specific gravity of PCC was tested with a water pycnometer according to the ASTM
D854-02. Beside, the oil absorption property of PCC was carried out according to the ASTM
D281-95.
X-ray Diffraction (XRD) analysis was performed using a Rigaku MiniFlex X-ray
Diffractometer (Rigaku Corporation, Tokyo, Japan) to investigate the characteristics of PCC.
This was measured over a range of 2θ from 10° to 70° and tested with CuKα (λ=1.5405Å) at
25°C in a reflection mode. The scanning rate and step size were 2°/min and 0.01°, respectively.
Tensile and flexural tests of manufactured composite samples were carried out using a
Model 5582 Advanced Mechanical Testing System (Instron Inc., Norwood, MA). Tensile
strengths were tested according to the ASTM D 638-99 and type І specimens of a dog-bone
shape were used. Flexural strengths were measured according to the ASTM D 790-03. Notched
izod impact strengths were measured with Tinius Olsen 92T Impact Tester (Tinius Olsen,
Horsham, PA) according to the ASTM D 256-05. Five replicate samples were used for each test.
For water absorption (WA) test, samples of a nominal size (30×12×3 mm) were prepared
from the injection molded plates and tested by measuring the weight of samples after 1- to 10-
11
week water immersion. The samples were removed from water bath at the end of each week. The
WA was calculated using:
WA(%)
Wa Wb
Wb
(2.1)
100
where Wa and Wb are sample weight (g) after and before soaking, respectively. Four
samples were used to get average value for each test.
2.3 RESULTS AND DISCUSSION
2.3.1 Basic Properties of Bamboo Filler
Figure 2.1(a) and Figure 2.1(b) show the thermo-gravimetric (TG) and derivative thermal
gravimetric (DTG) curves for RBF and GBP. The DTG curves of the two fibers represent
different trends. RBF shows one cellulose peak at 340°C with a broad tail of lignin. On the other
hand, GBP shows a main cellulose peak at 350°C and one shoulder of hemicellulose at 300°C
with a tail of lignin.(Brebu and Vasile 2010; Yao et al. 2008) To obtain the amount of each
component in bamboo fibers, deconvolutions for DTG curves of RBF and GBP were carried out
by using Gaussian fitting technique. (Hughes and Sexton 1988)
Table 2.1 shows the components and their percentage amount, acquired from the
deconvoluted DTG curves of GBP and RBF in the 150-600°C temperature range. It can be seen
that no change in holocellulose-1 (cellulose) contents was observed, but there were some
changes in lignin and holocellulose-2 (hemicellulose) contents between RBF and GBP. These
changes are related to the fiber disc refining process with steam at high temperature and
pressure.(Dwianto et al. 1996; Okubo, Fujii, and Yamamoto 2004; Salmen, Yin, and Berglund
2011; Winandy, Stark, and Horn 2008) Dwianto et al.(Dwianto et al. 1996) showed the initiation
of hemicellulose degradation in lignocellulosic fibers from 150oC and subsequent increases of
cellulose crystallinity.
12
100
o
Deriv. Weight (%/ C)
Weight (%)
80
1.0
(a)
RBF
GBP
60
40
20
RBF
GBP
(b)
Cellulose
0.8
0.6
Hemicellulose
0.4
Lignin
0.2
0.0
0
200
300
400
500
200
600
300
Temperature ( C)
600
1.0
(c)
o
No Silane
With Silane
0.8
Deriv. Weight (%/ C)
o
Deriv. Weight (%/ C)
500
Temperature ( C)
1.0
0.6
0.4
0.2
0.8
(d)
No Silane
With Silane
0.6
0.4
0.2
RBF
GBP
0.0
150
400
o
o
0.0
200
250
300
350
400
450
150
200
250
300
350
400
450
o
o
Temperature ( C)
Temperature ( C)
Figure 2.1 Thermogravimetric data of bamboo filler. (a) and (b): TG and DTG curves of
raw RBF and GBP; (c) and (d): DTG comparison between RBF and GBP before and after
silane treatment.
Table 2.1 Chemical compositions estimated from deconvoluted DTG curves for both GBP
and RBF
Components (%)
Fiber type
Total (%)
Hollocellulose-1 Hollocellulose-2
Lignin
Ash
GBP
28.71
25.97
15.03
20.29
100
RBF
28.81
27.87
14.21
19.11
100
13
Also, Y Yin et al.(Salmen, Yin, and Berglund 2011) suggested that the decreases of
hemicelluloses after steaming process were observed at 160oC and above. Considering the
published researches, the deconvoluted peak of hemicellulose in RBF might include some
portion of cellulose. Comparison of deconvoluted peak temperatures of GBP and RBF (273.69°C
and 277.67°C for holocellulose-2; 316.64°C and 309.11°C for holocellulose-1, respectively)
allowed the determination of the cellulose portion, which is included in holocellulose-2 of RBF.
The increase of holocellulose content after refining process played a significant role in the
thermal stability of bamboo fibers treated with silane coupling agent. As shown in Figure 2.1(c)
and Figure 2.1(d), the increase of DTG peak temperature is bigger from RBF to silane treated
RBF (TRBF) (from 339.56 to 358.22°C) than from GBP to silane treated GBP (TGBP) (from
342.69 to 350.86°C). This result indicates that the refining process in RBF helped develop more
silane crosslinks on the fiber surfaces.
Measured FTIR data is shown in Figure 2.2(a). The peak at 1739 cm-1 is assigned to the
O=C–OH group of the glucuronic acid unit and the decrease of this peak indicates a split-off of
the carbonyl group in hemicelluloses after steam assisted refining.(Salmen, Yin, and Berglund
2011; Sgriccia, Hawley, and Misra 2008) The comparison between RBF and GBP shows that
small amount of hemicelluloses was decomposed in RBF. Among GBP, TGBP and RBF, any
further remarkable differences were not shown. However, TRBF showed more critical
differences. The increased intensities of the peaks at 1604 cm-1 and at 1510 cm-1 indicate the
existence of silane crosslinks on the surface of TRBF after silane treatment and these changes
probably resulted from the traces of NH2 in silane coupling agent (Belgacem et al. 2004).
14
GBP
TGBP
O=C-OH
NH2
(a)
2000
RBF
1800
1600
Si-O-C
Si-O-Si
1400
1200
TRBF
1000
-1
Wave number (cm )
Figure 2.2 FT-IR spectra (a) of RBF and GBP before and after silane treatment and
schematic diagram (b) of bamboo fiber surfaces chemically modified by silane treatment.
15
Besides, the small peak shown around 1203 cm-1, arranged as Si-O-C or Si-O-Si band
was associated with silane crosslinks.(Lee et al. 2009) Above results suggest that the silane
coupling agent more effectively interacted with RBF than with GBP.
Figure 2.2(b) shows the schematic diagram transforming the hydrophilic bamboo fiber
surface to hydrophobic surface by silane coupling reaction. Through silanization, hydrolysis of
silane, condensation between silanols, and bond formation between siloxane and bamboo fibers
were consecutively achieved.(Lee et al. 2009)
2.3.2 Basic Properties of PCC
Figure 2.3(a) shows a SEM micrograph of PCC particles at the dry state and the averaged
individual particle size estimated from SEM micrograph amounts to 1.16μm. However, as shown
in this figure, individual cubic PCC particles were agglomerated together and these agglomerated
forms of PCC with measured mean particle size of 8.75 m were detected by the particle size
analysis using laser diffraction technique in Figure 2.3(b).
Besides, surface area, specific gravity and oil absorption of the PCC were 8.685 m2/g,
2.56 g/cm3, and 19.8 g/100g, respectively. The property ranges of commercial CC, commonly
used in plastics or WPCs are 2-24 m2/g, 2.7–2.9 g/cm3 and 13-21 g/100g for surface area,
specific gravity and oil absorption, respectively.(Wypych 2010b) Hence, the relatively large
surface area of PCC was related with smaller particle size and the somewhat reduced specific
gravity of PCC presumably resulted from impurities.
The XRD patterns of PCC are given in Figure 2.4(a). Through the comparison of XRD
peaks of the PCC samples with the Joint Committee on Powder Diffraction Standards (JCPDS)
data for three phases (calcite, aragonite and vaterite) of CC,(Tai and Chen 2008) it can be seen
16
that most PCC was calcite, which is thermally stable compared with other CC phases including
aragonite and vaterite. Also, the clear peaks of calcite indicate relative purity of the PCC.
Thermo-gravimetric analysis (TGA) data of the PCC is shown in Figure 2.4(b). It can be
seen that a small amount of weight loss (%) occurred around 280°C and a drastic loss happened
between 600 and 740°C. The loss at the higher temperature range was presumably weight loss of
CO2 (CaCO3 → CaO + CO2). The theoretical value of CO2 loss is 44 % for 100 % pure CC.
From the TGA data, the weight loss of CO2 was 41.99 %. Therefore, the calculated purity of
PCC (purity = measured weight loss / 44 × 100) was 95.43%.(Dweck 2010) The impurity was
organic matter from sugar processing. Derivative thermo-gravimetric (DTG) curve of PCC
further indicates the weight loss range of the PCC. The first low and broad peak at 283.40°C was
due to the decomposition of impurities included in PCC. The second high and sharp peak at
734.14°C was the mass loss of CO2 gas from decomposition of PCC.
2.3.3 Mechanical and Morphological Properties of Bamboo-PP/PE Composites
Figure 2.5 shows the flexural strength and modulus of GBP and RBF filled composites
with and without silane treatment.
For flexural strength (Figure 2.5(a)), GBP composite (GBP-C) showed a higher value
than that of RBF composite (RBF-C). However, after silane treatment, the strength of TRBF
composite (TRBF-C) became much higher than that of TGBP composite (TGBP-C). This result
shows that after silane treatment, less lignin and more holocellulose contents of RBF induced
more silane crosslinks,(Winandy, Stark, and Horn 2008) which led to the increase of flexural
strength through the enhanced adhesion and compatibility with polymer matrix.
17
15
D10=2.58µm
12
80
Volume (%)
D50=7.66µm
D90=15.25µm
9
DM=8.75µm
60
Span Factor=1.65
6
40
3
20
0
Cumulative Distribution (%)
100
(b)
0
0.1
1
10
100
Particle Size ( m)
Figure 2.3 Raw PCC morphology and particle size data. (a): SEM micrograph of PCC
particles; (b): measured particle size distribution based on the laser diffraction technique.
18
29.34
Intensity (a.u)
(a)
47.34 48.45
39.39
43.18
35.95
23.02
15
20
25
57.48
60.74
56.62
30
35
2
40
45
50
55
60
65
degrees)
100
8
CaCO3
80
CaO(s) + CO2(g)
6
Weight Loss (%)
Weight loss = 44%
60
4
Purity degree = 41.99 / 44 x 100% = 95.43%
40
2
20
0
Deriv. Weight (%/min)
(b)
0
0
100
200
300
400
500
600
700
800
900
o
Temperature ( C)
Figure 2.4 Measured XRD pattern (a) and thermal decomposition data (b) of the PCC filler.
19
35
(a)
Flexural Strength (MPa)
R-PP/PE : Fiber = 60 : 40
30
25
20
R-PP/PE
GBP-C
TGBP-C
RBF-C
TRBF-C
3
(b)
Flexural Modulus (GPa)
R-PP/PE : Fiber = 60 : 40
2
1
0
R-PP/PE
GBP-C
TGBP-C
RBF-C
TRBF-C
Figure 2.5 Comparison of measured flexural strength (a) and modulus (b) for R-PP/PE
resin and bamboo-PCC-PP/PE composites.
20
Figure 2.6 SEM micrographs of the impact fractured surfaces of TGBP (a) and TRBF (b)
composites.
21
For flexural modulus (Figure 2.5(b)), the difference between GBP-C and RBF-C seemed
to be due to the holocellulose content in bamboo fiber and an obvious effect of silane treatment
was not observed.
Also, the increased adhesion after silane treatment was verified from SEM images of
TGBP composite (TGBP-C) and TRBF composite (TRBF-C) as shown in Figure 2.6. Interfacial
gaps between fibers and polymer are clearly seen in TGBP-C (Figure 2.6(a)). On the other hand,
enhanced interfacial adhesions are observed in TRBF-C (Figure 2.6(b)).
Moreover, from these micrographs, the different fiber surfaces between RBF and GBP
were also seen. The surface of TGBP was relatively smooth with no individual fibers. TRBF had
a rough surface showing exposed individual fibers or fiber bundles. Similar results were
observed in the comparison of SEM micrographs between dry ground and steam exploded rice
husks (Donaldson, Wong, and Mackie 1988; Park et al. 2004).
According to Stark and Rowlands,(Stark and Rowlands 2003) the steam-aided disc
refining system led to the increase of fiber surface area and aspect ratio. Thus, it can be
explained that the significantly improved adhesion and compatibility between TRBF and plastic
matrix was resulted from the increased silane cross-links on its exposed fibers with the increase
of fiber surface area and aspect ratio.
2.3.4 Mechanical and Morphological Properties of PCC-PP/PE Composites
Figure 2.7 shows measured flexural, tensile and impact properties of the composites as a
function of PCC content. The flexural strength (Figure 2.7(a)) varied little from 0 to 10 wt%
PCC contents, but increased somewhat at the 15 wt% PCC content level. The highest increase
was shown at the 30% PCC content. The graph of flexural modulus showed similar features. The
tensile strength (Figure 2.7(b)) slightly decreased as PCC content increased and the tensile
22
modulus showed a slight increase. The notched impact strength (Figure 2.7(c)) decreased as a
function of PCC content (a 63% reduction at the 30% PCC level). The above results indicate that
the use of PCC led to decreased toughness for the blends of R-PP/PE with PCC.
Figure 2.8 shows the impact fractured surfaces of the blends of R-PP/PE with PCC. The
distribution and aggregation of PCC in the polymer matrix can be observed from Figure 2.8. The
neat R-PP/PE (Figure 2.8(a)) demonstrated clearly fractured features full of fibrils. The R-PP/PE
blended with 7 wt% PCC (Figure 2.8(b)) had some well dispersed PCC particles in the polymer
matrix (estimated mean particle diameter = 1.25 ± 0.43 µm). For the R-PP/PE blended with 30
wt% PCC (Figure 2.8(c)), some aggregated PCC particles were seen in the polymer matrix
(estimated mean particle diameter = 1.66 ± 1.24 µm). The data suggested that high shear forces
during compounding helped separate the agglomerated PCC particles (Figure 2.3). At higher
PCC loading levels (e.g., 30 wt%), some agglomeration still existed, leading to larger observed
particles. Overall, the separated PCC particles may still be too large to positively affect
toughness of the resultant composites.(Bryant and Wiebking 2002)
The observed mechanical properties of PCC-filled composite are in an argument of
general trend for mineral-filled plastic composites. For example, it was reported that the use of
talc, CC and Biodac® (a mineral-filled cellulosic material) in low density polyethylene led to a
significant increase in tensile modulus (200%, 77% and 130% for talc, CC and Biodac ®,
respectively) and decrease in toughness (42%, 58% and 86% for talc, CC and Biodac®,
respectively) at the 30% loading level.(Klyosov 2007)
2.3.5 Properties of PCC-Bamboo-PP/PE Composites
The mechanical properties of PCC-bamboo-PP/PE composites are summarized in Table
2.2. For the flexural strength of composites, the increased extents were not significantly different
23
until the 12 wt% PCC content level. However, a somewhat increased feature was observed at the
18 wt% PCC content. The flexural strengths of GBP filled composites with PCC were slightly
higher compared to those of RBF filled composites with PCC. After silane treatment, there was
little change for the flexural strengths of GBP filled composites with PCC regardless of PCC
content. This probably means that more lignin and less holocellulose contents in GBP with
relatively low fiber aspect ratio led to the decrease of silane crosslinks. (Gregorova et al. 2009;
Stark and Rowlands 2003) For RBF filled composites with PCC, over 15% increases of flexural
strengths were observed after the silane treatment in the whole range of PCC content, indicating
that less lignin and more holocellulose contents in RBF led to the increase of flexural strengths
by improved silane coupling.(Gregorova et al. 2009; Inoue and Norimoto 1991; Stark and
Rowlands 2003) On the other hand, impact strengths decreased as PCC content increased as a
result of the early fractures from the lack of toughening effect due to the size of PCC.
Impact fractured SEM images of GBP filled composites with 6 and 18 wt% PCC contents
are shown in Figure 2.9. In the 6 wt% PCC reinforced composite (Figure 2.9(a)), PCC particles
were evenly dispersed. However, the 18 wt% PCC reinforced composite (Figure 2.9(b))
demonstrated frequent advents of PCC aggregations with the poor interfacial adhesions between
fiber and polymer matrix.
The enhanced moduli of variously treated bamboo filled composites as a function of PCC
content are shown in Figure 2.10. The flexural and tensile moduli of GBP and RBF filled
composites with PCC had similar values from 12 to 18 wt% PCC contents. However, the values
of TGBP and TRBF filled composites with PCC featured the continuous increases. The different
trends of measured modulus among the above composites are attributed to the aggregation and
dispersion of PCC in bamboo-plastic matrix.
24
For GBP and RBF filled composites with PCC, the small changes of flexural and tensile
modulus from 12 to 18 wt% PCC contents indicate that the aggregated PCC particles prevented
from shortening the ligament thickness and led to a faster initiation of fracture (Fu, Wang, and
Shen 1993; Fu and Wang 1993; Thio et al. 2002).
Table 2.2 Mechanical properties of PCC-bamboo-PP/PE composites
System
a
R-PP&PE/GBP
R-P&PE/TGBP
R-PP&PE/RBF
R-PP&PE/TRBF
a
PCC (wt%)
0
6
12
18
0
6
12
18
0
6
12
18
0
6
12
18
Tensile
modulusb
(GPa)
2.43 (0.19)B
2.39 (0.15)B
2.84 (0.17)A
2.88 (0.13)A
2.58 (0.06)D
2.77 (0.28)C
2.83 (0.08)B
3.18 (0.19)A
2.63 (0.24)B
2.88 (0.25)B
3.51 (0.24)A
3.64 (0.13)A
2.75 (0.24)C
2.74 (0.15)C
3.21 (0.07)B
3.48 (0.18)A
Flexural
Flexural
modulus
strength
(GPa)
(MPa)
1.98 (0.05)B 29.84 (0.82)B
1.94 (0.04)B 29.79 (0.38)B
2.42 (0.03)A 30.52 (0.61)B
2.44 (0.04)A 31.40 (0.55)A
2.01 (0.05)D 30.45 (0.67)A
2.19 (0.04)C 30.96 (0.29)A
2.52 (0.07)B 30.94 (0.67)A
2.82 (0.06)A 30.90 (0.38)A
2.33 (0.03)C 28.07 (0.46)C
2.52 (0.07)B 28.49 (0.27)BC
2.95 (0.05)A 29.69 (0.37)AB
2.98 (0.17)A 29.17 (0.99)A
2.32 (0.05)D 33.44 (0.31)B
2.57 (0.03)C 33.60 (0.25)B
2.85 (0.04)B 33.89 (0.33)B
3.12 (0.17)A 34.58 (0.81)A
a
Impact
strength
(kJ/m2)
4.33 (0.15)A
4.10 (0.17)B
3.63 (0.10)C
3.40 (0.20)D
4.29 (0.24)A
3.69 (0.27)B
3.36 (0.26)C
3.22 (0.15)C
3.74 (0.17)A
3.11 (0.16)B
2.69 (0.08)C
2.38 (0.05)D
3.15 (0.06)A
2.87 (0.10)B
2.82 (0.05)B
2.71 (0.02)C
R-PP&PE: recycled polypropylene and polyethylene; GBP: ground bamboo particle; TGBP:
silane treated bamboo particle; RBF: refined bamboo fiber; TRBF: silane treated bamboo fiber.
R-PP&PE to bamboo fiber ratio was 60 to 40 wt%. PCC was based on total composite weight.
b
Means with the same letter for each category are not significantly different at the 95%
confidence level; numbers in parenthesis are standard deviations.
25
1.2
(a)
Strength
Modulus
1.0
30
0.8
25
0.6
20
Flexural Modulus (GPa)
Flexural Strength (MPa)
35
0.4
0
10
20
30
PCC (wt%)
1.2
(b)
Strength
Modulus
1.0
25
0.8
20
0.6
15
Tensile Modulus (GPa)
Tensile Strength (MPa)
30
0.4
0
10
20
30
PCC (wt%)
(c)
2
Impact Strength (kJ/m )
6
5
4
3
2
0
10
20
30
PCC (wt%)
Figure 2.7 Comparison of flexural (a), tensile (b) and impact (c) properties of PCC-filled
composites as a function of PCC content.
26
Figure 2.8 SEM micrographs of the impact fractured surfaces of R-PP/PE and PCC blends.
(a): R-PP/PE; (b): R-PP/PE and PCC (7 wt%); (c): R-PP/PE and PCC (30 wt%).
27
Figure 2.9 SEM micrographs of the impact fractured surfaces of PCC-GBP-PP/PE
composites: (a) 6 wt% PCC and (b) 18 wt% PCC.
28
4
Flexural Modulus (GPa)
(a)
3
2
GBP-C
TGBP-C
RBF-C
TRBF-C
1
0
6
12
18
PCC (wt%)
4.0
Tensile Modulus (GPa)
(b)
3.5
3.0
GBP-C
TGBP-C
RBF-C
TRBF-C
2.5
2.0
0
6
12
18
PCC (wt%)
Figure 2.10 Comparison of flexural modulus (a) and tensile modulus (b) of PCC-GBP and
PCC-RBF composites before and after silane treatment as a function of PCC content.
29
For TGBP and TRBF filled composites with PCC, the remaining silanols, which were not
cross-linked with bamboo fibers might react with the surface of PCC, leading to the improved
dispersion and decreased aggregation of PCC in composites and the large modulus changes from
12 to 18 wt% PCC contents.
Table 2.3 shows measured water absorption (WA) values as a function of time among
different composites with and without PCC. The addition of PCC (18 wt%) into bamboo-plastic
matrix gave rise to a slight increase of WA properties. The WA values of composites with PCC
were about 1-2% higher compared to those of composites without PCC after 10 weeks water
immersion. It is noticed that WA values in this experiment were from 2.37-3.27% at the 18 wt%
PCC level after a two-week water immersion, while other previous research data for the WA
values of CC filled WPC showed 7.1-9.0% at 10-20% CC levels (Klyosov 2007). This
discrepancy is presumably due to the differences in manufacturing methods, base resins,
composite density, and fibers (wood versus bamboo). The WA increases of PCC filled
composites seemed to be related to mainly the interfacial gap from poor compatibility between
hydrophilic PCC and polymer. Among variously treated bamboo filled composites, it is observed
that TRBF filled composites had the lowest WA values, while RBF filled composites had the
highest values. This fact indicates that the moisture uptakes in composites are related to the WA
properties of respective fibers and compatibility between fibers and plastic matrix.
Two-way ANOVA test data on WA in the variation of immersion time are shown in
Table 2.4. The ANOVA results showed that both fiber type and PCC content had a significant
influence on WA of the composites. For the combined effect of fiber type and PCC content on
WA, a significant interaction was observed after 1-week water immersion at the 95% confidence
30
level. However, after that, interactions between fiber type and PCC content on WA were not
significant.
Table 2.3 WA (%) of variously treated bamboo-PP/PE composites with and without PCC.
PCC (wt%)
WA (%) at time intervala
a
Fiber type
0
18
1 week
2 weeks
3 weeks
4 weeks
10 weeks
GBP
2.13 (0.28)
3.01 (0.54)
3.78 (0.74)
4.45 (0.95)
6.42 (1.34)
RBF
2.02 (0.14)
3.07 (0.33)
4.03 (0.43)
4.83 (0.50)
7.30 (0.52)
TGBP
1.83 (0.26)
2.42 (0.44)
3.06 (0.56)
3.52 (0.69)
5.52 (1.24)
TRBF
1.32 (0.02)
2.03 (0.20)
2.41 (0.04)
2.87 (0.09)
4.57 (0.31)
GBP
2.37 (0.44)
3.27 (0.86)
4.24 (1.21)
5.57 (0.97)
7.81 (1.05)
RBF
2.92 (0.31)
4.47 (0.66)
5.86 (0.79)
7.02 (0.85)
8.91 (0.06)
TGBP
1.89 (0.26)
2.87 (0.45)
3.88 (0.33)
4.52 (0.55)
6.72 (0.69)
TRBF
1.63 (0.14)
2.37 (0.28)
3.10 (0.38)
3.78 (0.48)
6.15 (0.77)
a
GBP: ground bamboo particle; RBF: refined bamboo fiber; TGBP: silane treated bamboo
particle; TRBF: silane trated bamboo fiber; the numbers in parenthesis are standard deviations
Table 2.4 Two-way ANOVA tests of the effect of fiber and PCC on WA of R-PP/PE
composites at different time intervals.
Water absorption at time intervala
Variables
1 week
2 weeks
3 weeks
4 weeks
10 weeks
Fiber
< 0.0001
< 0.0001
< 0.0001
< 0.0001
< 0.0001
PCC
0.0004
0.0024
0.0004
< 0.0001
< 0.0001
Fiber × PCC
0.0197
0.1174
0.1876
0.2434
0.9606
a
The values shown are p-values of 2-way ANOVA tests. The p-values smaller than 0.05 indicate
significant influences of the corresponding treatments on WA at the 95% confidence level.
31
2.4 CONCLUSIONS
The differences in chemical composition and surface morphology between RBF and GBP
led to more effective silane crosslinking on the RBF surfaces. Dry PCC particles showed an
agglomerated form, and compounding PCC with R-PP/PE resin helped separate it to smaller
individual particles. Measured flexural strength and flexural modulus of PCC-only-filled
composites increased significantly from 15 to 30% PCC content levels, while the tensile and
impact strength of composites decreased with the addition of PCC.
For composites with hybrid bamboo and PCC fillers, tensile and flexural moduli were
improved with the increase of PCC content. After silane treatment of the bamboo filler, RBF
filled composites showed noticeably increased mechanical properties compared to those of GBP
filled composites. For modulus values, PCC-bamboo-polymer composites were 3-4 times higher
than those of PCC-polymer composites at high PCC levels.
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35
CHAPTER 3 MECHANICAL AND PHYSICAL PROPERTIES OF CORE-SHELL
STRUCTURED WOOD PLASTIC COMPOSITES: EFFECT OF SHELLS WITH
HYBRID MINERAL AND WOOD FILLERS
3.1 INTRODUCTION
Coextrusion generally consists of two or more extruders combined with one die to
produce multiple-layer products (Rosato 1998). By combining molten multiple plastic layers
with various properties into one profile, optimization of product performance is possible (Giles,
Wagner, and Mount 2005). Water resistance, air entrapment, oxygen barrier, and increased
toughness are some of the advantages of coextruded products (Giles, Wagner, and Mount 2005;
Rosato 1998). For example, coextrusion method has been applied to produce packaging films for
maintaining content freshness, pipes requiring high mechanical properties, and other specialty
products (Doshi, Charrier, and Dealy 1988; Kim et al. 1984).
Coextrusion in wood plastic composite (WPC) was first reported with a combination of
WPC core and pure plastic shell layer. Stark and Matuana (Stark and Matuana 2007) investigated
moisture uptake, flexural properties and weathering performance between non-coextruded and
coextruded WPC with a pure high density polyethylene (HDPE) or pure polypropylene (PP)
shell. In their research, coextruded WPC demonstrated much reduced moisture uptake than noncoextruded WPC and there were almost no differences in flexural properties between them. They
further studied the effects of a stabilized shell layer blended with HDPE and additives including
a compatibilizer, a photostabilizer and a nanosized titanium dioxide (TiO2) on the coextruded
WPC after weathering tests (Stark and Matuana 2009). The coextruded WPC with a stabilized
shell layer showed better water resistance properties and the combination of a compatibilizer and
a photostabilizer improved composite lightening by a synergistic effect. The individual use of
TiO2 or a photostabilizer showed a noticeably enhanced color stability for each capped WPC.
36
However, a simultaneous use of both additives had a deleterious effect on color stability. Jin and
Matuana (Jin and Matuana 2008) studied the improvements of water resistance through the
application of a pure HDPE plastic shell onto an extruded WPC core. The water absorption (WA)
and thickness swelling (TS) properties of coextruded WPC were better than those of noncoextruded WPC without a shell.
The application of a pure plastic shell with a relatively low modulus over a WPC core
negatively affected overall composite modulus. Jin and Matuana (Jin and Matuana 2010)
reported a study using a PVC shell layer on a PVC-wood flour (WF) core. Increased flexural
strength and decreased flexural modulus were observed for the coextruded composites. It was
pointed out that high strength and low stiffness characteristics of PVC led to the decrease of
composite flexural modulus. To acquire a simultaneous enhancement of both flexural modulus
and flexural strength, they used carbon nano-tube (CNT) in a shell layer, in combination with
two core compositions (40% versus 60% WF) and two processing conditions (low versus high
temperature). The combination of a CNT-filled PVC shell and a high temperature processed
WPC core showed better flexural strength and flexural modulus. However, the use of high-cost
CNTs was not cost-effective for WPC products. Yao and Wu (Yao and Wu 2010) investigated
the coextruded WPC using the combinations of recycled PE core compositions (weak and strong)
and a virgin PE shell with different WF contents and shell thicknesses. In the comparison of
mechanical properties between weak and strong core systems, coextruded composites with a
weak core showed much higher percentage increases in flexural and impact strength than those
with a strong core. At the same shell thickness, less WF loading in the shell led to highly
increased impact strength, but there were almost no changes in flexural modulus. On the other
hand, the increase of impact strength and the decrease of flexural modulus were observed with
37
increased shell thickness. Thus, they suggested that good flexural modulus of coextruded WPC
resulted from the combination of a thinner shell and a higher modulus core.
The addition of mineral fillers into polymeric materials enhances their physical and
mechanical properties (Nwabunma and Kyu 2008; Wypych 2010a). Among the mineral fillers,
precipitated calcium carbonate (PCC) from chemical processes (Bleakley and Jones 1993; Liu et
al. 2006; Porter and Wilson 1999; Rothon 1999; Wypych 2010a) has characteristics of very fine
and regular particle size (Xanthos 2005). Thus, its uses in plastic composites have led to
improved mechanical properties (Patnaik 2003; Zuiderduin et al. 2003). However, the fine size
of PCC conversely causes its aggregations in composites, leading to a deleterious effect on the
properties of PCC filled composites (Lam et al. 2009). Hence, to decrease this deleterious effect,
surface treated PCC has been used. Lam et al. (Lam et al. 2009) investigated the effect of
modified PCC in PP matrix through optical and thermal analysis. Nanosized PCC (n-PCC) and
nanosized surface modified PCC (ns-PCC) filled PP composites were prepared. The initial
decomposition temperature of ns-PCC/PP composites was higher than that of n-PCC/PP
composites at the same filling ratio, indicating that the reinforcing effect of ns-PCC was
increased by surface treatment. J. Cayer-Barrioz et al. (Cayer-Barrioz et al. 2006) studied the
interfacial adhesion between polyamide 66 (PA66) and PCC (untreated PCC; 3wt% stearic acid
treated PCC; amino acid treated PCC). Dynamic mechanical analysis (DMA) demonstrated the
effects of the surface treated PCC through the G” peak shifts by +13K for the amino acid
treatment and +6K for the stearic acid treatment. Moreover, environmental scanning electron
microscope analysis confirmed the differences of adhesions among PCC filled composites. Kim
et al. (Kim et al. 2012) investigated that technical feasibility of using PCC of sugar origin as a
reinforcing filler for bamboo fiber plastic composite. It was shown that air-dry PCC particles
38
were in an agglomerated form made of individual cubic particles of about 1.2 micron in diameter.
Compounding with plastic resin helped separate the PCC to smaller individual particles. The use
of PCC led to a significant increase of flexural strength and flexural modulus of PCC filled
composites after 10 wt% PCC loading level. For bamboo filled composites with PCC, tensile
modulus and flexural modulus were improved with the increase of PCC content and with use of
surface treated bamboo filler. The use of raw and treated PCC in coextruded WPC (e.g., in the
shell layer to modify overall composite properties) has not been reported.
The objectives of this study were to elucidate the effect of TPCC and WF loadings in the
shell layer on the mechanical, water absorption and thermal expansion properties of coextruded
WPC with two different core systems.
3.2 EXPERIMENTAL
3.2.1 Materials and Preparation
Fluffy form of recycled-PP/HDPE resin from a local plastic recycling company was
pelletized using a twin-screw extruder before blending with other raw materials. Virgin HDPE
(HGB 0760) was provided by ExxonMobile Chemical Co. (Houston, TX, USA). Pine WF (20
mesh particle size) was supplied by American Wood Fibers Inc. (Schofield, WI, USA). PCC was
obtained from Domino Sugar Co. (New Orleans, LA, USA) and screened to pass an 100-mesh
screen. MAPE (EpoleneTM G2608) from Eastman Chemical Co. (Madison, TN, USA) and silane
coupling agent (Z-6094) from the DOW Corning Co. (Midland, MI, USA) were utilized to
increase the compatibility between fillers and plastic matrix. Lubricant (TPW 306) from Struktol
Co. (Stow, OH, USA) was used to improve the processing of WPC profile.
The silane treatment of PCC was made by an immersion method. Before treatment, PCC
was dried at 80°C for 24 hr. After that, aqueous solution (ethanol:water = 90:10) including silane
39
coupling agents (1 wt%) was prepared and left at 25°C for 30 min to attain proper hydrolysis and
condensation of silane. The PCC was then immersed in the solution to achieve proper silane
surface treatments. Silane treated PCC was dried at 80°C for 24 hr again. The recycled PP/HDPE
plastic and virgin HDPE were used as base resins for two core systems (i.e., weak core and
strong core). Virgin HDPE was used to make shell material for both systems. Materials used for
shell were compounded and pelletized at 165 (feeder), 175, 180, 170 and 180oC (die) before the
coextrusion process. TPCC was first compounded with virgin HDPE to form a 50/50 wt% master
batch. The compounded master batch pellets were then compounded again with WF at target
loading levels.
3.2.2 Coextruded WPC Manufacturing
The composites were formulated with two core types (i.e., weak and strong) in
combination with two shell systems for each core type. The formulation for the weak- and
strong-core systems were, respectively, R-PP/HDPE: WF: Lubricant: MAPE = 40: 50: 6: 4 wt%
and V-HDPE: WF: Lubricant: MAPE = 40: 50: 6: 4 wt%. The two shell systems were a) TPCC
(6, 12, and 18%) + WF (15%) + V-HDPE (79, 73, and 67%) and b) TPCC (12%) + WF (0, 5, 15,
and 25%) + V-HDPE (88, 83, 73, and 63%).
The composites were manufactured with a pilot-scale coextrusion system (Yao and Wu
2010). This system consists of a Leistritz Micro-27 co-rotating parallel twin-screw extruder
(Leistritz Corporation, Allendale, NJ) for core and a Brabender 32 mm conical twin-screw
extruder (Brabender Instruments Inc., South Hackensack, NJ) for shell. A specially-designed die
with a cross-section area of 13 x 50 mm and a target shell thickness of 1.0 mm was used. A
vacuum sizer was used to maintain the targeted size. The coextruded profiles passed through a 2
m water bath with water spraying using a down-stream puller. Manufacturing temperatures for
40
core were controlled between 150 and 175oC. Manufacturing temperature for shells varied from
150 to 170oC in the variation of different shell formulations.
3.2.3 Characterization
X-ray photoelectron spectroscopy (XPS) was carried out using a Kratos Axis-165 high
performance multi analysis (Kratos Analytical Ltd., Manchester, UK) to investigate atomic
elements of treated PCC with a monochromatic MgKa source (1253.6 eV) at 15 kV and 20 mA.
Oven dried PCC particles were mounted onto a holder and placed in a vacuum of 1–5 x 10-8 torr.
FT-IR analysis was performed using a Bruker Tensor 27 system (Bruker Optics Corporation,
Billerica, MA) to investigate the changes the compositions of PCC before and after silane
treatment. Powder samples were directly placed on the top sample plate with the ZnSe crystal
illuminated and then analyzed in the range of 525–4000 cm-1. The acquired FT-IR data were the
averages of 32 scans.
Three-point flexural test was conducted using a model 5582 Instron testing machine
(Instron Co, Norwood, MA) at a crosshead speed of 6mm/min according to the ASTM D790.
Tinius Olsen Mode 1892 impact tester (Tinius Olsen Inc., Horsham, PA) was used to test Izod
impact strength without notching according to the ASTM D256, using samples with a 3-mm
thickness. These samples were acquired by cross-cutting the extruded profiles such that the
impact force on the test samples was perpendicular to the coextrusion direction. Five samples
were used for each group. Type of impact facture failure (i.e., hinge versus complete) for each
sample was recorded.
Water soaking properties including WA and TS were measured using samples of 254 ×
50 × 12 mm in size according to the ASTM D7031-2004. WA and TS values were determined as:
41
WA(%)
TS (%)
Wa Wb
Wb
Ta
Tb
(3.1)
100
(3.2)
100
Tb
Where, Wa and Wb are sample weights (g) after and before soaking; Ta and Tb are sample
thickness (mm) after and before soaking. Three samples per group were used to get an average
value for each test.
For the coefficient of thermal expansion (CTE) measurement of coextruded WPCs,
samples of 63.5 × 50 × 12 mm in size were prepared and tested. Length measurements were
done with an IDF-130E Mitutoyo digimatic indicator (Mitutoyo Corporation, Japen) of 0.01 mm
accuracy along the long (i.e., extrusion) dimension of the extruding direction of respective
samples (between the two opposing center points on the end surfaces). Samples were conditioned
to reach equilibrium temperatures of-15oC and then 60oC prior to measurements at each
temperature. CTE was calculated as:
CTE ( )
Lf
Li
Li
(T f
(3.3)
Ti )
Where, Lf and Li are the final and initial measured sample lengths (mm); Tf and Ti are the
final and initial measured temperature (oC). Three samples were used to get an average CTE
value for each group.
3.3 RESULTS AND DISCUSSION
3.3.1 Properties of Treated PCC
Figure 3.1 shows measured XPS spectra of untreated PCC (Figure 3.1(a)) and TPCC
(Figure 3.1(b)).
42
(a)
Intensity (Counts/sec)
O 1s
Ca 2p
C 1s
Ca 2s
N 1s
O 2s
Si 2s
600
500
400
300
200
Si 2p
100
0
Binding Energy (eV)
Intensity (Counts/sec)
(b)
O 1s
C 1s
Ca 2p
Ca 2s N 1s
Si 2s
600
500
400
300
200
Si 2p
100
O 2s
0
Binding Energy (eV)
Figure 3.1 XPS spectra of untreated PCC (a) and treated PCC (b).
43
Table 3.1 Element composition, oxygen-carbon, and silicon-oxygen ratios of untreated PCC
and TPCC from the XPS measurement.
C1s
Element composition (%)
O1s
Ca2p
Si2p
N1s
Element ratio
O/C
Si/O
Untreated PCC
48.92
38.66
8.76
2.11
1.54
0.79
0.05
TPCC
58.25
28.73
5.25
3.55
4.22
0.49
0.12
Type
Before silane treatment, the magnitudes of silicon and nitrogen peaks in PCC were small.
After treatment with the silane coupling agent (z-6094), these peak magnitudes were
significantly increased. To acquire quantitative information of the element composition of PCC,
the peak area of each electron from XPS spectra was standardized with a sensitivity factor.
The ratios of element composition (O/C and Si/O) resulted from the standardized process
are shown in Table 3.1. Regardless of treatment conditions, the order of percentage composition
was carbon (largest), oxygen, and calcium (smallest). However, for silicon and nitrogen, the
order of percentage composition changed before and after the treatments. The silane treatment
led to an increase of C1s electron from 48.92% to 58.25% and a decrease of O1s electron from
38.66% to 28.73%. The lower O/C and the higher Si/O ratios of silane treated PCC compared to
untreated PCC demonstrated silane coupling agent attachment to the surface of PCC. There were
174 % increase of N1s and 68 % increase of Si2p electron, which showed the existence of silane
element on treated PCC.
The silane deposition onto PCC is also observed in FT-IR spectra (Figure 3.2). For TPCC,
the peak at 1040 cm-1 is assigned to Si-O-Si linkages considered as siloxan on the PCC. Beside,
the peaks at 1100 cm-1 and 1080 cm-1 are associated with longer Si–O-Si structures
(Abdelmouleh et al. 2004).
44
Transmittance (%)
(a)
Region A
PCC
TPCC
1600
1500
1400
1300
1200
1100
1000
900
-1
Wave number (cm )
Region A
Transmittnace (%)
(b)
Si-O-Si
Si-O-Si
1150
1125
1100
1075
1050
1025
-1
Wave number (cm )
Figure 3.2 FT-IR spectra of PCC and TPCC (a: whole spectra and b: selected region).
45
Figure 3.3 Schematic diagram of PCC surface chemically modified by silane treatment.
46
Figure 3.3 shows the schematic diagram illustrating the transformation of the hydrophilic
PCC surface to hydrophobic surface by silane coupling reaction. Through silanization,
hydrolysis of silane, condensation between silanols, and bond formation between siloxane and
PCC surface were achieved in a consecutive manner.
3.3.2 Mechanical and Morphological Properties of Composites
Table 3.2 lists summarized data of measured flexural strength and modulus of coextruded
WPCs with two core systems and various TPCC and WF contents in a shell layer. Figure 3.4 (a)
and (b) show comparative charts of the two properties for various systems.
The flexural modulus for weak and strong cores was similar (2.75 GPa vs. 2.79 GPa).
However, the flexural strength for the strong core was much higher than that of the weak core
(34.5 MPa vs. 14.9 MPa). This was due to the fact that the strong core was made of virgin HDPE,
while weak core was made of the recycled PP/HDPE.
For the composite flexural strength, the weak core system showed significant increases in
relation to the strength value of weak core only. In this system, the shell layer made of filled
virgin HDPE effectively prevented the deformation and subsequent crack occurrence of the
composite due to its higher tensile strength. However, composites with the strong core system
showed decreased overall strength compared with the corresponding core-only control. In this
system, the core-only control had a relatively high strength. Hence, the interfaces between shell
and core layers made of the same virgin HDPE and different filler contents played a dominant
role in controlling flexural strengths regardless of TPCC loading.
For flexural modulus, both core systems had a decreased moduli compared to their
respective core only controls. The extent of decrease was bigger in the weak core system than
that in the strong core system. This is probably associated with the relatively poor compatibility
47
between shell and core plastics in the weak core system (Giles, Wagner, and Mount 2005;
Rosato 1998). Beside, the somewhat enhanced flexural strength and modulus with the increase of
TPCC contents presumably reflected the toughening effect of TPCC.
Table 3.2 Mechanical and thermal expansion properties of coextruded core-shell WPCs
with various levels of TPCC and WF in the shell layer.
System
a
TPCC
contents in
WF shell
(weak core)
TPCC
contents in
WF shell
(strong core)
WF contents
in TPCC shell
(weak core)
WF contents
in TPCC shell
(strong core)
PCC or
wood in
shell (wt%)
No shell
6
12
18
No shell
6
12
18
No shell
0
5
15
25
No shell
0
5
15
25
Flexural
Flexural
b
modulus
strength
(GPa)
(MPa)
2.75 (0.10)A
14.90 (1.02)B
1.81 (0.18)B
20.57 (0.60)A
1.95 (0.14)B
20.83 (1.30)A
2.00 (0.14)B
21.34 (0.79)A
2.79 (0.06)A
34.50 (0.59)A
2.50 (0.23)B
30.11 (1.62)B
2.54 (0.11)AB 30.93 (1.56)B
2.62 (0.14)AB 30.51 (0.98)B
2.75 (0.10)A
14.90 (1.02)C
1.72 (0.10)C
19.49 (0.61)A
1.78 (0.15)BC 19.69 (1.18)A
1.95 (0.14)B
20.83 (1.30)A
1.81 (0.20)BC 17.74 (2.01)B
2.79 (0.06)A
34.50 (0.59)A
2.31 (0.13)C
28.97 (1.25)C
2.34 (0.09)C 29.34(0.81)BC
2.54 (0.11)B
30.93 (1.56)B
2.21 (0.06)C
26.43 (0.40)D
a
Impact
strength
(kJ/m2)
2.42 (0.26)B
5.30 (0.98)A
5.01 (0.93)A
4.86 (0.47)A
6.05 (0.75)B
6.94 (1.11)A
6.64 (1.01)A
5.96 (0.73)B
2.42 (0.26)E
7.40 (1.65)A
6.23 (1.25)B
5.01 (0.93)C
3.61 (0.90)D
6.05 (0.75)B
8.17 (1.05)A
7.72 (1.04)A
6.64 (1.01)B
5.26 (0.98)C
CTE
(×10-5/ oC )
6.29 (0.03)A
6.35 (0.97)A
5.94 (0.89)A
6.01 (0.58)A
4.37 (0.69)B
6.13 (0.17)A
5.32 (0.61)AB
5.61 (0.63)A
6.29 (0.03)AB
7.32 (0.61)A
5.92 (1.10)B
5.94 (0.89)B
5.28 (0.78)B
4.37 (0.69)B
6.14 (0.89)A
6.24 (0.44)A
5.32 (0.61)AB
5.22 (0.73)AB
Weak Core: R-PP/PE:WF:Lubricant:MAPE = 40:50:6:4; Strong Core: V-HDPE:WF:Lubricant:
MAPE = 40:50:6:4
b
Means with the same letter for each category are not significantly different at the 95%
confidence level; numbers in parenthesis are standard deviations.
48
Bar: strength
Line: modulus
Flexural Strength (MPa)
(a) Weak Core
4
30
3
20
2
10
1
0
Flexural Modulus (GPa)
40
0
Core only
6
12
18
TPCC in Shell (wt.%)
4
Bar: strength
Line: modulus
Flexural Strength (MPa)
(b) Strong Core
30
3
20
2
10
1
0
Flexural Modulus (GPa)
40
0
Core only
6
12
18
TPCC in Shell (wt.%)
Weak Core: Right
Strong Core: Left
(c)
2
Impact Strength (kJ/m )
10
8
6
4
2
0
Core only
6
12
18
TPCC in Shell (wt.%)
Figure 3.4 Flexural (a: weak core, b: strong core) and impact (c) properties of coextruded
core-shell structured WPCs-effect of TPCC content in a 15% WF-filled shell.
49
4
Bar: strength
Line: modulus
Flexural Strength (MPa)
(a) Weak Core
30
3
20
2
10
1
0
Flexural Modulus (GPa)
40
0
Core only
0
5
15
25
WF in Shell (wt.%)
4
Bar: strength
Line: modulus
Flexural Strength (MPa)
(b)Strong Core
30
3
20
2
10
1
0
Flexural Modulus (GPa)
40
0
Core only
0
5
15
25
WF in Shell (wt.%)
10
Weak Core: Right
Strong Core: Left
2
Impact Strength (kJ/m )
(c)
8
6
4
2
0
Core only
0
5
15
25
WF in Shell (wt.%)
Figure 3.5 Flexural (a: weak core, b: strong core) and impact (c) properties of coextruded
core-shell structured WPCs-effect of WF content in a 12% TPCC-filled shell.
50
Impact strengths of the coextruded WPCs are shown in Table 3.2 and Figure 3.4(c). As
shown, the strength improvements were, respectively, 119% and 15% for weak and strong core
systems. The more flexible and tougher shell layer made of virgin HDPE formed a more
effective protection for the relatively brittle core layer in the weak core system. Thus, in
coextruded WPC manufacturing, the combination of low-quality recycled plastic core and high
quality virgin plastic shell can be a cost-effective choice with acceptable mechanical properties.
Flexural strengths for the coextruded WPCs with two core systems and various WF
contents in a 12% TPCC filled shell are shown in Table 3.2 and Figure 3.5(a) and (b).
In the weak core system, the flexural strengths of coextruded WPCs were significantly
higher than that of the core-only control. For the strong core system, coextruded WPCs showed
lower strengths than that of their core-only control. Both coextruded WPC systems showed the
slightly increased flexural strengths with increased WF content in the shell up to the 15% level.
A similar trend was observed in flexural modulus of both coextruded systems. However,
excessive WF loading in the shell layer (e.g., 25%) resulted in the reduction of flexural
properties due to the poor WF dispersion and subsequent stress concentration. Impact strengths
for both coextruded WPC systems are shown in Table 3.2 and Figure 3.5(c). A similar large
increase in the impact strength of both core systems was observed in comparison with their
respective core-only controls. The coextruded WPCs with TPCC only in shell showed the
highest impact strength values among the coextruded WPCs. This is mainly related to the
toughening effect of TPCC in a shell layer. However, as WF was added to the shell layer, the
impact strengths of both systems were decreased, indicating that the addition of WF led to early
initiation of cracks and failures in the shell layer due to its relatively bigger size and consequent
cavity.
51
Figure 3.6 SEM micrographs of impact fractured surfaces of coextruded core-shell
structured WPC: (a) weak core with TPCC-only filled shell layer; (b) weak core with
TPCC and WF filled shell layer.
52
As a result, in the weak core system, the impact strength at the 25% WF filled shell
showed the lowest value. Therefore, to acquire higher impact strength values in coextruded
WPCs, the amount and size of WF in a shell layer should be adjusted and the combined use of
smaller sized TPCC particle is encouraged.
Figure 3.6 shows SEM micrographs with magnified core-shell boundaries of the weakcore coextruded WPCs. Noticeable boundaries were observed in both SEM micrographs due to
composition difference between WF-filled R-PP/PE core and TPCC- or TPCC/WF-filled HDPE
shell.
The fractured surface of the coextruded WPC with TPCC-only filled shell (Figure 3.6(a))
showed clearly debonded small holes from the toughening effect of TPCC. On the other hand,
the fractured surface of coextruded WPC with TPCC and 15% WF filled shell (Figure 3.6(b))
showed the irregularly debonded large holes with fracture propagation traces from the fast and
extensive failures in the shell matrix indicating the weakened shell with large wood fiber
presence.
3.3.3 Impact Fracture Type Comparison
Figure 3.7 shows a comparison of fracture type distributions (i.e., hinge versus complete)
from impact tests of core-shell structured WPCs.
As shown, the fracture types varied with core quality (weak versus strong), composite
structure, and filler compositions in the shell layer.
In the weak core systems (Figure 3.7(a) and (b)), while core-only controls had 100%
complete-fracture mode, coextruded WPCs had much decreased complete-fracture percentage
(conversely, much increased hinge fracture). The percentages of hinge fractures were 95, 70, and
53
65% at the TPCC levels of 6, 12, and 18% in a 12% WF filled shell layer; and 95, 90, 70, and 60%
at the WF levels of 0, 5, 15, and 25% in a 15% TPCC filled shell layer.
For the strong core systems (Figure 3.7(c) and (d)), while core-only control had about 90%
complete-fracture mode, coextruded WPCs had somewhat decreased complete fracture
percentage. The percentages of hinge fractures were 40, 35, and 30% at the TPCC levels of 6, 12,
and 18% in a 12% WF filled shell layer; and 40, 35, 35, and 30% at the WF levels of 0, 5, 15,
and 25% in a 15% TPCC filled shell layer.
The observed impact fracture process in core-shell structured WPCs is systematically
illustrated in Figure 3.8.
As shown, after an incident force impacted the front shell layer of coextruded WPCs, the
shell layer absorbed some of the impact energy. The remaining impact force, which passed
through the front shell layer, was absorbed by the core layer, and was lastly absorbed by the back
shell layer.
In the weak core systems, when the impact force hit the shell surface of the coextruded
WPC, more flexible and tougher shell layer made of virgin HDPE formed an effective protection
for the relatively brittle core layer. However, the relatively poor compatibility between core and
shell led to the unstable interfacial adhesion, which diverged the impact force.
The impact force was further distributed in the core layer and the force in the direction of
loading was decreased to such an extent that it could not cause complete fracture of the back
shell layer. While a core-only control showed a straight complete-fracture mode, coextruded
WPCs had a much curved fracture line with hinged back shell from diverged impact forces in a
core layer. Some apparent separation showing core-shell de-bonding was observed (Figure
3.8(a)). On the other hand, for the strong-core systems, the core and shell layers formed a
54
stronger bond resulting from the same HDPE resin and acted a whole system to resist the impact
force. There was little diversion of the incident impact force, which caused complete fracture in
the direction of impact loading (Figure 3.8(b)).
Complete
Hinge
100
80
80
Percentage (%)
Percentage (%)
Hinge
100
60
40
60
40
20
20
(a)
(b)
0
0
Core only
6
12
Core only
18
TPCC in Shell (wt.%)
Hinge
0
5
15
25
WF in Shell (wt.%)
Complete
Hinge
100
100
80
80
Percentage (%)
Percentage (%)
Complete
60
40
Complete
60
40
20
20
(d)
(c)
0
0
Core only
6
12
Core only
18
0
5
15
25
WF in Shell (wt.%)
TPCC in Shell (wt.%)
Figure 3.7 Fracture type distributions (hinge versus complete) from impact tests of
coextruded core-shell structured WPCs. (a) and (b): weak core; (c) and (d): strong core.
55
Figure 3.8 Schematic diagram showing impact fracture process of coextruded core-shell
structured WPCs and photographs of typical impact fractured samples. (a): weak core
system; (b): strong core system.
56
3.3.4 WA and TS Properties of Composites
The relationship between WA and TS for weak core and strong core systems in the
variation of TPCC and WF content in a shell layer during 54 days of water immersion is shown
in Figure 3.9.
For both core systems, it can be seen that the shell layer of coextruded WPCs prevented
the core layer from excessive moisture uptake. Moreover, the comparison of WA and TS values
between weak and strong core coextruded WPCs shows that the strong-core system had much
better WA properties. This was due to the good bonding between WF and virgin HDPE in the
strong core compared with a relatively poor bonding among different recycled materials in the
weak core.
Increased loading of TPCC in the shell layer did not lead to any remarkable increases in
WA and TS values of coextruded WPCs regardless of core qualities, as shown in Figure 3.9(a)
and (b). This result is considerably ascribed to the improved dispersibility of TPCC and
increased compatibility between TPCC and V-HDPE matrix in the shell layer. The relationship
between WA and TS for weak- and strong-core systems in the variation of WF content in a shell
layer is also shown in Figure 3.9(c) and (d). The water-proof effects of a shell layer were
observed in both systems with the increased WF content in the shell layer. However, at the 25%
WF level, WA and TS percentages of both systems were significantly increased. This was
presumably related to the more WF loading and its hydrophilic property. Also, the high WF
content in the shell layer induced cavities or gaps between WF and plastic matrix from the more
aggregations of WF, consequently, leading to higher WA and TS values.
57
8
Weak Core (No Shell)
6% TPCC in Shell
12% TPCC in Shell
18% TPCC in Shell
6
Thickness Swelling (%)
Thickness Swelling (%)
8
4
2
Weak Core (No Shell)
No WF in Shell
5% WF in Shell
15% WF in Shell
25% WF in Shell
6
4
2
(a)
(c)
0
0
0
2
4
6
8
10
12
0
2
8
10
12
2.5
Strong Core (No Shell)
6% TPCC in Shell
12% TPCC in Shell
18% TPCC in Shell
Thickness Swelling (%)
Thickness Swelling (%)
6
Water Absorption (%)
Water Absorption (%)
2.5
2.0
4
1.5
1.0
0.5
2.0
Strong Core (No Shell)
No WF in Shell
5% WF in Shell
15% WF in Shell
25% WF in Shell
1.5
1.0
0.5
(b)
0.0
0.0
0.5
1.0
1.5
2.0
2.5
3.0
(d)
3.5
0.0
0.0
Water Absorption (%)
0.5
1.0
1.5
2.0
2.5
3.0
3.5
Water Absorption (%)
Figure 3.9 TS and WA properties of coextruded core-shell WPCs. (a): weak core and (b)
strong core- effect of TPCC loading in a 15% WF-filled shell. (c): weak core and (d) strong
core- effect of WF loading in a 12% TPCC-filled shell.
58
3.3.5 Thermal Expansion Properties of Composites
Table 3.2 lists the CTE values for coextruded WPCs with two different core systems in
the variation of WF and PCC loading in a shell layer. For core-only controls, the weak core
system had higher CTE values compared with the strong core system (6.2×10-5/oC vs. 4.2×105 o
/ C) due to poor bonding between recycled plastics and wood fibers in the weak core system.
Coextruded composites with 12% TPCC only in shell or the combination of 15% WF and 6%
TPCC in shell showed increased CTE values. However, as filler amount in the shell layer
increased, the CTE values of coextruded WPCs were reduced. Thus, a small amount of filler was
not enough to offset the large CTE of the plastic matrix. As a result, coextruded WPCs with a
larger portion of plastic in the shell layer would have increased CTE values.
The weak core system with recycled plastic core and virgin plastic shell at high filler
loading in a shell layer showed smaller CTE values compared with core-only controls. Thus, a
reinforced shell layer can help control overall thermal expansion properties of coextruded
composites with a relatively poor core system. It seems that both wood and PCC fillers in the
shell led to similar CTE values of the overall composites. From this result, it is found that the
dimensional changes of the coextruded WPC can be significantly affected by filler loading and
plastic quality in a shell layer.
3.4 CONCLUSIONS
Significant differences existed in mechanical, WA/TS and CTE properties between
coextruded WPCs with weak and strong core systems as influenced by the shell compositions. In
the weak core system with R-PP/HDPE, coextruded composites with a reinforced shell showed
significantly improved flexural strengths compared to their core-only composite. In the strong
core system made of V-HDPE, the flexural strengths of coextruded WPC were lowered
59
compared with the core-only composite. Impact strengths were significantly improved for both
coextruded systems at low shell filling levels. High level of filler loading in the shell led to some
decreases of impact strength in both coextruded systems. The types of impact fracture (i.e., hinge
versus complete) varied largely with core quality and filler composition in the shell. WA/TS
values of coextruded WPCs with high TPCC content in a shell layer were lower than those of
core-only composites and coextruded WPCs with high wood content in the shell layer. The use
of high percentage of plastic in a shell layer led to increased overall CTE values for coextruded
WPCs. However, increased filler loading in the shell layer led to the decrease of CTE values for
resultant coextruded WPC, especially, for the weak core system. Thus, the combination of a
relatively weak core (e.g., made of recycled plastics) and a reinforced shell (e.g., with small
sized inorganic particles) could be a cost-effective system for coextruded WPC manufacturing.
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Stark, N.M., and L.M. Matuana. 2009. Co-extrusion of WPCs with a clear cap layer to improve
color stability. In 4th Wood Fiber Polymer Composites International Symposium, 1-13.
Bordeaux, France.
Wypych, G. 2010. Handbook of fillers, 3th ed: ChemTec Publishing.
Xanthos, M. 2005. Functional fillers for plastics: Wiley-Vch Co.
Yao, F., and Q.L. Wu. 2010. Coextruded Polyethylene and Wood-Flour Composite: Effect of
Shell Thickness, Wood Loading, and Core Quality. Journal of Applied Polymer Science 118:
3594-601.
Zuiderduin, W.C.J., C. Westzaan, J. Huetink, and R.J. Gaymans. 2003. Toughening of
polypropylene with calcium carbonate particles. Polymer 44: 261-75.
61
CHAPTER 4 MECHANICAL PROPERTIES OF CO-EXTRUDED WOOD PLASTIC
COMPOSITES WITH GLASS FIBER FILLED SHELL
4.1 INTRODUCTION
As a new generation green composite, coextruded wood plastic composite (WPC) with a
core-shell structure has been recently developed and used to enhance performance characteristics
of WPC (Jin and Matuana 2010; Stark and Matuana 2007). By proper combination of
constituting layers one can achieve a balance of such properties as light weight, high strength,
high stiffness, wear resistance, biological resistance, unusual thermal expansion characteristics,
appearance, etc (Jin and Matuana 2010; Kim and Wu 2012; Stark and Matuana 2007; Yao and
Wu 2010). Fundamental understanding of the interactions between shell and core layer with
different structural (e.g., thickness) and material combinations is, however, needed to achieve
desired product performance.
Early published work with WPC co-extrusion was done with unfilled plastic shell layers.
Stark and Matuana (Stark and Matuana 2007) studied the moisture uptake of non-coextruded
WPC and coextruded WPC with a high density polyethylene (HDPE) shell layer. It was shown
that coextruded WPCs exhibited much reduced moisture absorptions than a non-coextruded
WPC, leading to much needed water resistance improvement. Jin and Matuana (Jin and Matuana
2010) investigated flexural properties of coextruded WPCs using a PVC shell layer and reported
that the addition of a pure plastic shell with a relatively low modulus over a WPC core
negatively affected overall composite modulus in a core-shell structured material. Thus, to get a
simultaneous improvement for both flexural modulus and flexural strength of coextruded WPCs,
they considered the incorporation of carbon nano-tube (CNT) in a PVC shell layer. The addition
of CNT filler in a shell layer led to some improvement of both flexural properties, but the use of
62
CNTs can lead to significant manufacturing cost increases. Hence, the attempts to improve the
mechanical properties of coextruded WPCs using cheaper and more abundant fillers in a shell
layer have been considered. Yao and Wu (Yao and Wu 2010) evaluated the coextruded WPC
using the combinations of recycled PE (LDPE and HDPE) core compositions and a virgin HDPE
shell with different shell thickness and wood fiber (WF) contents in a shell layer. In their
experiment, coextruded WPCs with a weak core having more LDPE showed higher percentage
improvements in flexural and impact strength compared with coextruded WPCs with strong core
having less LDPE. At the same shell thickness, more WF contents in the shell led to the decrease
of impact strength with no changes in flexural modulus. As the shell thickness increased, the
increase of impact strength and the decrease of flexural modulus were observed. As a result, it
was suggested that the better flexural modulus of coextruded WPC can be acquired from the
combination of a thinner shell and a higher modulus core. More recently, Kim et al. (Kim and
Wu 2012) studied the effects of silane treated precipitated calcium carbonate (TPCC) of sugar
origin and WF filled shell layers on the properties of coextruded WPC with two different core
systems (recycled PP/PE and virgin HDPE core). In this study, weak-core coextruded WPCs
using a recycled PP/PE (over 90% of PP) core and a virgin HDPE shell showed significantly
improved flexural strengths compared with core-only WPC regardless of filler content in a shell
layer. However, strong-core coextruded WPCs made of the same virgin HDPE core and shell
demonstrated relatively low strength increases in relation to core-only WPC and even showed
strength reductions at high filler content in a shell layer. In impact strength, weak-core
coextruded WPCs with different core-shell thermoplastics showed a 206% improvement at
TPCC only in a shell layer. The incorporation of fillers in a shell layer of coextruded WPCs has
somewhat positively affected the mechanical properties of core-shell structured composites.
63
However, cost and performance effective shells are still needed to achieve desired composite
performance.
Short glass fibers (SGFs) have been used as a good reinforcing filler for WPC. Jiang et al.
(Jiang et al. 2003) reported that impact strength of WPC using PVC and WF was considerably
increased by adding SGF (a 60 % improvement in notched impact strength with 5 % SGF filled
WPC). This improvement was due to the 3-dimensional network structure formed by the WF and
SGF. It was suggested that the effectiveness of SGFs reinforcement was highly associated with
SGF aspect ratio, SGF/WF composition, and PVC content. Jiang et al. (Jiang et al. 2007) applied
a commingled unidirectional GF-polymer composite sheet (CUGPCS) to surface of WPC deck
profile and achieved a good interfacial adhesion between CUGPC and WPC. Significantly
increased bending modulus of rupture (MOR), modulus of elasticity (MOE), and strain at break
were reported for the laminated WPC composites. However, manufacturing CUGPCS reinforced
WPC with premade WPC boards requires additional processing steps and energy and long-term
effectiveness of bonded surfaces has not been demonstrated. On the other hand, coextrusion
technology making various properties of WPC highly tunable through fully capped shell in one
step could be effective to achieve synergetic effects of SGF and WPC.
The objective of this study was to investigate the effects of various GF contents in a shell
layer and shell thickness changes on the flexural and impact properties of coextruded WPC in
combination with three core systems (low, moderate, and strong).
4.2 EXPERIMENTAL
4.2.1 Materials and Preparation
HDPE (AD60-007) and LDPE (LD103) were provided by ExxonMobile Chemical Co.
(Houston, TX, USA). Pine WF (20 mesh particle size) was supplied by American Wood Fibers
64
Inc. (Schofield, WI, USA). Short GF reinforced HDPE pellets were provided by RTP Co.
(Winona, MN, USA). The material was of the type RTP 707 CC UV Natural with glass fiber
content of 40 % by the weight of total formulation. The fiber diameter was 0.014 mm and fiber
length was 4 mm prior to compounding. The fibers were sized with a silane–based solution
before compounding. UV (Type 6022) and coupling (Fusabond EMB 100D) agents were added
to the GF-HDPE system during compounding. MAPE (EpoleneTM G2608) from Eastman
Chemical Co. (Madison, TN, USA) was utilized to increase the compatibility between wood
fillers and plastic matrix. Lubricant (TPW 306) from Struktol Co. (Stow, OH, USA) was also
used to improve the processing of WPC profile. The mixtures of HDPE, LDPE and GF filled
HDPE were used as base resins to make three core systems (i.e., weak, moderate and strong).
HDPE and 40% GF filled HDPE or their mixtures were used to make a shell material as a
function of GF content. Materials used for shell were used without pre-compounding to prevent
further breakage of GF fibers.
4.2.2 Injection Molding of GF-HDPE Samples
Injection molding was used to make test samples for establishing shell properties.
Experimental design included five blends with GF/HDPE ratios of 0:100, 10:90, 20:80, 30:70,
and 40:60. Melt compounding was first performed using an intermesh, counter-rotating
Brabender twin-screw extruder (Brabender Instruments, Hackensack, NJ) with a screw speed of
40 rpm. The temperature profile of barrels ranged from 150 to 175°C. The extrudates were aircooled and then pelletized into granules. The granules were injection-molded into standard
mechanical test specimens using a Batenfeld Plus 35 injection molding machine (Batenfeld, NJ,
USA). The injection temperatures were 190 and 180 °C for HDPE/GF composites and virgin
HDPE, respectively. All specimens were then conditioned for 72 hr at a temperature of 25 °C and
65
a relative humidity of 50% for later characterization.
4.2.3 Co-extruded WPC Manufacturing
The composites were formulated with three core types (i.e., weak, moderate and strong)
in combination with different GF contents in shell and shell thickness changes for each core type.
The formulations for the weak-, moderate- and strong-core systems were, respectively, LDPE:
HDPE: WF: Lubricant: MAPE = 30: 10: 50: 6: 4 %, LDPE: HDPE: WF: Lubricant: MAPE = 10:
30: 50: 6: 4 %, and HDPE: GF: WF: Lubricant: MAPE = 50: 20: 25: 3: 2 %. Six GF content
levels (i.e., 0, 10, 15, 20, 30, and 40 %) and five shell thicknesses (i.e., 0.8, 1.0, 1.2, 1.4, and 1.6
mm) were used to make different shells.
The composites were manufactured with a pilot-scale coextrusion system (Yao and Wu
2010). This system consists of a Leistritz Micro-27 co-rotating parallel twin-screw extruder
(Leistritz Corporation, Allendale, NJ) for core and a Brabender 32 mm conical twin-screw
extruder (Brabender Instruments Inc., South Hackensack, NJ) for shell. A specially-designed die
with a cross-section area of 13 × 50 mm was used. A vacuum sizer was used to maintain the
targeted size. The coextruded profiles passed through a 2 m water bath with water spraying using
a down-stream puller. Manufacturing temperatures for core were controlled at 155 (feeder), 160,
165, 170, 170, 170, 160, 150, 140, 130, and 155oC (die). Manufacturing temperature for shells
varied from 150 to 165oC in the variation of different shell formulations. Different shell
thicknesses were achieved by controlling shell feeding rate and extrusion speed.
4.2.4 Mechanical Testing and Characterization
Three-point flexural test was conducted using a model 5582 Instron testing machine
(Instron Co, Norwood, MA) at a crosshead speed of 6 mm/min according to the ASTM D790 to
determine bending strength and modulus for both injection molded and extruded samples. Tinius
66
Olsen Mode 1892 impact tester (Tinius Olsen Inc., Horsham, PA) was used to test Izod impact
strength according to the ASTM D256. The injection molded samples were notched prior to
testing. For extruded material, test samples of 3-mm in thickness were acquired by cross-cutting
the extruded profiles such that the impact force on the test samples was perpendicular to the
coextrusion direction. These samples were tested without notching to preserve original core-shell
structure. More than 5 samples were tested for each group. Type of impact facture failure (i.e.,
hinge versus complete) for each sample was recorded.
Composite morphologies were characterized by a Quanta 3D FEG Dual Beam Scanning
Electron Microscope (SEM) with Focused Ion Beam (FIB) (FEI Company, Hillsboro, OR). The
samples were coated with Pt to improve the surface conductivity before observation and
observed at an acceleration voltage of 5 kV.
4.3 RESULTS AND DISCUSSION
4.3.1 Shell GF-HDPE Composite
4.3.1.1 Morphology
Typical SEM micrographs for 30% GF-filled HDPE composites are shown in Figure 4.1.
Most GF fillers were aligned perpendicular to the fracture plane (i.e., along the injection molding
flow direction). Certain filler pullouts happened during fracture process as indicated by the
circular voids on the fracture plane. The remaining GF fillers appeared to be well-bonded to the
matrix, indicating effect of surface coupling treatments of the GFs prior to mixing with the
plastic matrix.
4.3.1.2 Mechanical Property
Flexural property (MOE and MOR) and impact strength of GF filled shell-only
composites are summarized in Table 4.1 and the data are plotted in Figure 4.2(a).
67
Flexural moduli of GF filled composites exhibited an increasing trend with increased
filler loadings. The neat HDPE showed a flexural modulus of 0.85 GPa and for 40% GF filled
composite, the modulus was increased to 5.8 GPa. The increase of flexural modulus was
attributed to the enhanced interfacial interaction between the GF filler and polymer matrix and
the high aspect ratios of GF fillers in the plastic matrix also helped it (DiBenedetto 2001; Rothon
1999; Wypych 2010a; Xanthos 2005).
In flexural strength, GF-filled composites showed significant strength increases with
increased GF loadings. At the 40% GF loading level, the strength was 3.94 times higher than that
of the neat resin. It can be concluded that GF filler alignments as shown in the SEM micrographs
(Figure 4.1) also played an important role in determining the flexural strength of the composites
(Fu and Lauke 1996)
Figure 4.2(b) illustrates the notched Izod impact strengths of GF-filled composites in the
variation of GF content.
The neat HDPE had an impact strength of 28.57 kJ/m2 and when the GF filler was added
to the system, the strength was significantly decreased. At the 10% GF loading level, the impact
strength was 9.63 kJ/m2 and the reduced impact strength of GF-filled composites may be due to
the decrease of ductility and toughness by the incorporation of fillers, restricting the molecular
chain movement of the plastic matrix (Nunez et al. 2003). In GF-filled composites, the impact
strength increased with further increase of GF loading beyond the 10% level. Since GFs were
well bonded to the plastic matrix, the perpendicularly aligned GF fillers to the impact testing
direction (Figure 4.1) could help absorb the impact force, leading to enhanced impact strength.
68
Figure 4.1 SEM micrographs of impact fractured surfaces of 30% GF filled HDPE
composites.
Table 4.1 Mechanical properties of core-only WPCs with different core quality (weak,
moderate, and strong) and shell-only composites with various GF content.
Composite
Type
Shell
Core
a
Material
Make-up
GF/HDPE: 0:100
GF/HDPE: 10:90
GF/HDPE: 20:80
GF/HDPE: 30:70
GF/HDPE: 40:60
Weak
Moderate
Strong
Property
MOEa (GPa)
0.85 (0.06)
1.31 (0.03)
2.34 (0.33)
3.57 (0.08)
5.78 (0.19)
2.16 (0.05)
3.23 (0.07)
4.45 (0.30)
The numbers in parenthesis are standard deviations.
69
MOR (MPa)
21.83 (0.96)
29.22 (0.18)
40.92 (0.58)
57.80 (0.69)
85.93 (1.59)
22.32 (0.54)
31.04 (0.41)
48.82 (1.52)
IM (kJ/m2)
28.57 (2.04)
9.62 (0.37)
10.37 (0.37)
11.94 (0.16)
14.55 (0.34)
5.39 (0.42)
3.52 (0.26)
4.60 (0.60)
7
100
MOE
MOR
80
5
4
60
3
40
MOR (MPa)
MOE (GPa)
6
2
20
1
(a)
0
0
0
10
20
30
40
GF Content (%)
2
Impact Strength (kJ/m )
35
30
25
20
15
10
(b)
5
0
0
10
20
30
40
GF Content (%)
Figure 4.2 Flexural properties (a) and impact strength (b) of GF filled shell-only composites.
70
4.3.2 Extruded Core-only Composites
4.3.2.1 Morphology
SEM micrographs from impact fractured core-only WPCs with three different core
systems (i.e., weak, moderate, and strong) are shown in Figure 4.3. For weak- and moderate-core
WPCs (shown in Figure 4.3(a) and (c)), well distributed WF fillers in plastic matrix are seen in
both composites. But, more magnified images (Figure 4.3(b) and (d)) show somewhat different
fracture features between the two composites. As shown in Figure 4.3(b), for weak-core WPC,
broken WF filler still covered with plastic indicates that more flexible LDPE in the matrix
prevented the WFs from being cleanly pulled-out. However, for moderate core WPC with HDPE
(Figure 4.3(d)), relatively clean surfaces of WFs without plastic covering are observed. For
strong-core WPC (Figure 4.3(e)), well-dispersed WF and GF fillers are apparent and the close-up
micrograph (Figure 4.3(f)) indicates good interfacial adhesion between fillers (both GF and WF)
and plastic matrix. Thus, it can be possibly said that more coupling agent (and lubricant)
included in the strong-core WPC effectively improved the dispersibility and adhesion of fillers in
plastic matrix (Tselios et al. 1999).
4.3.2.2 Mechanical Property
Figure 4.4(a) and (b) summarize measured flexural properties (MOE and MOR) and
impact property of core-only WPCs with three different core systems (i.e., weak, moderate, and
strong, respectively). HDPE has more rigid and tougher characteristics compared with LDPE due
to its macromolecular properties (Charles 2000). Hence, in WPC matrix containing more HDPE,
better flexural modulus and strength are expected. The MOE and MOR values were increased in
the order of weak, moderate, and strong core.
71
Figure 4.3 SEM micrographs of core-only WPCs: weak core (a and b), moderate core (c
and d) and strong core (e and f).
72
60
5
(a)
MOE
50
MOR
40
3
30
2
20
1
MOR (MPa)
MOE (GPa)
4
10
0
0
Weak
Moderate
Strong
Core Quality
Complete
Hinge
100
80
6
60
4
40
2
Percentage (%)
(b)
2
Impact Strength (kJ/m )
8
20
0
0
Weak
Moderate
Strong
Core Quality
Figure 4.4 Flexural properties (a) and impact strength with percentage of breaking types (b)
for extruded core-only WPCs (weak, moderate, and strong core).
73
The weak- and moderate-core WPCs included 50% WF, while the strong-core WPC had
25% WF and 20% GF. The MOE and MOR values of core-only composites are significantly
higher in the strong-core system than those in weak- and moderate-core systems. This indicates
that the incorporation of GF in WPC matrix led to the remarkably improved flexural properties
due to its high fiber strength and high aspect ratios (Ghaus and Hamid 2008; Shakeri and
Raghimi 2010; Thwe and Liao 2002; Tungjitpornkull, Chaochanchaikul, and Sombatsompop
2007; Velmurugan and Manikandan 2007). Moreover, the GF fiber direction (perpendicular to
flexural test direction) in extruded composites affected the strength improvements of the strong
core WPC (Fu and Lauke 1996).
Impact strength values were from high to low in the order of weak, strong, and moderate
core systems. This result seems to indicate that weak-core WPC with more flexible LDPE (75%
LDPE in plastic) absorbed more impact energy than strong-core WPC and moderate-core WPC
using HDPE-only and less LDPE (25% LDPE in plastic), respectively.
As shown in Figure 4.4(b), the impact fracture types varied with different core quality.
Weak core WPC showed 90% hinge-fracture mode. On the other hand, strong- and moderatecore WPCs showed the decreased hinge-fracture mode (40 % and 10 %, respectively).
4.3.3 Coextruded Core-Shell Composites
4.3.3.1 Morphology
Figure 4.5 shows SEM micrographs for the cross sections (Figure 4.5(a) and (b)) and the
laterally-cut shell surface (Figure 4.5(c) and (d)) of strong-core coextruded WPC with 40% GFfilled shell. More magnified image (Figure 4.5(b)) of cross section shows more frequent
existence of GF fillers in shell (40% GF) region than that in core (20% GF) region. When it
74
comes to filler alignment in the shell layer, the alignment of GF fillers in polymer matrix is the
same as coextrusion profile direction, which is apparently shown in Figure 4.5(d).
Figure 4.5 SEM micrographs for cross section (a and b) and top view (c and d) of strongcore coextruded WPC with 40% GF shell.
75
Figure 4.6 SEM micrographs of coextruded WPCs: moderate core with PE shell (a) and
with 40% GF shell (b); strong core with PE shell (c) and with 40% GF shell (d).
Impact fractured morphologies for moderate- and strong-core coextruded composites are
shown in Figure 4.6. Moderate core with WF filler (left) and HDPE-only shell (right) are clearly
seen in Figure 4.6(a) and also, moderate core with WF filler (left) and 40% GF-filled shell (right)
are observed in Figure 4.6(b).
76
As shown in the right section of Figure 4.6(b), the direction of GF filler alignments in the
40% GF-filled shell layer is parallel to that of impact testing direction and this means that it is
hard to expect an improved impact property in coextruded composites by the addition of GF
filler into the shell layer. Strong core with WF/GF-filled shell (right) and HDPE-only shell (left)
are differentiated from each other in Figure 4.6(c) and strong core with WF/GF-filled shell (left)
and 40% GF-filled shell (right) in Figure 4.6(d) are discerned by the existence of WF. Though
some interfacial gaps between core and shell layers were observed, their bonding condition looks
reliable.
4.3.3.2 Mechanical Property
Measured mechanical properties (MOE, MOR, and impact strength) for coextruded
composites with three different core qualities are summarized in Table 4.2. Summary plots of the
data as a function of the corresponding shell and core properties are shown in Figures 4.8, 4.9,
and 4.10, respectively for MOE, MOR, and impact strength. Also, in the figures, the plots of allcore and all-shell properties are illustrated as comparisons.
For each mechanical property, the two lines of all-core and all-shell composites formed
upper and lower bonds, respectively. The core property was constant for each coextruded
composite system, while the shell property increased with the increase of GF content in the shell
layer. The intersections represented the points where core and shell had the same property values.
When the coextruded composite systems changed from weak to moderate and strong cores, the
intersection points shifted toward higher property value points.
77
Table 4.2 Mechanical properties of coextruded composites with different core quality (weak, moderate, and strong) and shell
thickness including various GF content in shell layer.
GF
contenta (%)
0
10
15
20
30
40
a
Shell
Thickness (mm)
0.8
1.0
1.2
1.4
1.6
0.8
1.0
1.2
1.4
1.6
0.8
1.0
1.2
1.4
1.6
0.8
1.0
1.2
1.4
1.6
0.8
1.0
1.2
1.4
1.6
0.8
1.0
1.2
1.4
1.6
Weak Core
MOE
(GPa)
1.88
1.77
1.66
1.60
1.56
2.16
2.07
1.96
2.11
2.08
2.18
2.31
2.40
2.46
2.51
2.38
2.55
2.66
2.77
2.70
2.94
3.13
3.34
3.55
3.83
3.46
3.44
3.70
3.95
4.66
MOR
(MPa)
21.42
21.52
21.25
20.60
20.99
25.68
25.49
25.36
25.54
26.89
23.79
25.87
28.75
29.81
31.42
24.68
27.62
30.55
32.46
32.79
31.21
34.48
36.62
38.65
39.34
34.70
35.65
36.49
39.71
42.49
Moderate Core
IM
(kJ/m2)
7.14
8.56
10.42
12.65
15.89
5.52
7.42
8.24
8.00
11.44
6.50
7.79
9.30
8.00
11.61
6.71
7.26
8.36
7.22
11.61
7.22
8.03
8.24
9.75
11.98
8.48
6.61
8.11
9.24
12.47
MOE
(GPa)
2.85
2.76
2.56
2.60
2.27
2.99
2.94
2.60
2.77
2.57
3.14
3.08
2.86
2.79
2.90
3.22
3.20
3.06
3.14
3.24
3.55
3.66
3.80
3.93
4.11
3.97
4.06
4.13
4.32
5.09
The filler content was based on the total composite weight.
78
MOR (MPa)
31.57
31.40
32.31
32.60
28.51
32.76
33.76
32.53
34.53
32.53
34.19
36.02
35.52
34.40
35.90
34.50
35.69
36.49
37.28
38.24
36.32
38.90
43.14
44.51
46.85
41.86
43.00
44.87
46.96
55.60
Strong Core
IM
(kJ/m2)
N.A.
7.26
9.27
N.A.
11.42
6.22
5.37
7.99
N.A.
7.59
4.77
5.31
5.71
N.A.
7.67
7.35
5.40
9.64
N.A.
7.74
5.18
5.33
5.83
N.A.
8.06
4.60
5.89
6.40
8.83
8.59
MOE
(GPa)
3.66
3.20
2.41
2.45
2.26
3.65
3.43
3.11
2.81
2.90
3.88
3.68
3.54
3.42
3.30
4.12
3.92
3.88
3.77
3.64
4.49
4.62
4.73
4.86
5.03
4.84
4.98
5.12
5.27
5.41
MOR
(MPa)
43.19
38.41
29.92
33.34
30.67
42.98
41.96
40.58
37.64
37.85
45.72
45.02
44.36
42.86
42.78
47.27
47.74
47.86
46.50
42.65
51.75
52.55
53.84
54.33
54.83
53.34
53.95
56.26
57.71
58.70
IM
(kJ/m2)
N.A.
6.73
9.25
10.92
13.51
6.78
8.04
8.56
10.50
11.74
7.61
8.29
9.20
10.22
11.78
7.08
8.29
9.21
9.35
11.81
N.A.
10.03
8.99
10.14
11.96
N.A.
8.35
8.94
10.32
12.20
For each composite system, composite modulus increased with the increase of shell
modulus at a given shell thickness. When the shell modulus was less than the core modulus, the
increase of shell thickness led to reduced composite modulus. For example, when the sell
thickness was doubled from 0.8 to 1.6 mm, modulus for composites with pure HDPE-shell was
reduced from 1.88 to 1.56 GPa for weak core, from 2.85 to 2.27 GPa for moderate core, and
from 3.66 to 2.26 GPa for strong core, respectively. Thus, thinner shell should be preferred for
this system. When the shell modulus was higher than the core modulus, the opposite is true. The
increase of shell thickness led to increased composite modulus. For example, when the sell
thickness was doubled from 0.8 to 1.6 mm, modulus for composites with 30% GF-filled shell
was increased from 2.94 to 3.83 GPa for weak core, from 3.55 to 4.11 GPa for moderate core,
and from 4.49 to 5.03 GPa for strong core, respectively. Thus, thick shell should be preferred for
this system.
In composite strengths, relatively small strength increases from 30% to 40% GF levels
are observed for each composite system. This is because excessive GF fillers in the plastic matrix
of the shell layer made GFs broken through fiber-fiber interaction and consequently, led to
relatively low strength increases (Fu et al. 1999; Gupta et al. 1989; Thomason 2002). Also, at the
thinnest shell thickness (i.e., 0.8 mm), composite strength increases stayed in low level and this
is attributed to the GF breakages on the surface of the shell layer resulting from the high shear
forces of coextrusion process (Truckenmuller and Fritz 1991).
A fiber shape filler is one of the primary load-carrying elements in composites and thus,
structural efficiencies of the fiber filled composites are highly dependent on the fiber alignments
in composites (Kacir, Narkis, and Ishai 1975). Typically, GF-filled extruded composite profiles
show high percentages of GF filler alignments for extruding direction and when a convergent die
79
is used, their experimental flexural moduli reach about 90% of the corresponding theoretical
maximums (Hine et al. 1997). Dominant GF filler alignments for coextruding direction are
observed in the laterally-cut shell surface (Figure 4.5(c)) of GF-filled coextruded composites and
therefore, it can be concluded that the GF alignments improved the flexural property of
coextruded composites (Figure 4.7 and Figure 4.8).
However, in impact strength, the GF alignments were ineffectively activated. The
comparison of coextruded composites with HDPE-only and GF-filled shells (Table 4.2 and
Figure 4.9) shows no strength improvements with the increase of GF loading in the shell layer.
4.3.4 Analysis of Composite Property
To statistically evaluate the reliability of the mechanical property (MOE, MOR, and
impact strength) improvements for core coextruded composites in the variation of GF content in
the shell layer and shell thickness, two-way analysis of variance (ANOVA) method was adopted
and tested. ANOVA test results are shown in Table 3 for weak-, moderate-, and strong-core
coextruded composites, respectfully. In weak-core coextruded composite, the ANOVA results
showed that both GF content and shell thickness had a significant influence on each mechanical
property of the composites and also, for the combined effect of GF content and shell thickness, a
significant interaction was observed for all mechanical properties. Moderated-core coextruded
composite also showed the ANOVA results having significantly different p-values for GF
content, shell thickness, and the interaction between GF content and shell thickness. But, the
effect of shell thickness for MOE shows 0.0447, which is close to 95% confidence level. In case
of strong-core coextruded composites, similar ANOVA test results were observed except for the
effect of GF content on impact strength not showing significant influences (P-value = 0.2531).
80
7
all core
(a)
Composite MOE (GPa)
6
all shell
shell=0.8mm
5
shell=1mm
shell=1.2mm
4
shell=1.4mm
shell=1.6mm
3
2
1
0
0
1
2
3
4
5
6
7
Shell MOE (GPa)
Composite MOE (GPa)
7
all core
(b)
6
all shell
shell=0.8mm
5
shell=1mm
shell=1.2mm
4
shell=1.4mm
shell=1.6mm
3
2
1
0
0
1
2
3
4
5
6
7
Shell MOE (GPa)
7
Composite MOE (GPa)
all core
(c)
6
all shell
shell=0.8mm
5
shell=1mm
shell=1.2mm
4
shell=1.4mm
shell=1.6mm
3
2
1
0
0
1
2
3
4
5
6
7
Shell MOE (GPa)
Figure 4.7 Comparison of MOE for coextruded composites with all core- and all shellcomposites: (a) weak core, (b) moderate core, and (c) strong core.
81
Composite MOR (MPa)
100
all core
(a)
80
all shell
shell=0.8mm
shell=1mm
shell=1.2mm
60
shell=1.4mm
shell=1.6mm
40
20
0
0
20
40
60
80
100
Shell MOR (MPa)
Composite MOR (MPa)
100
all core
(b)
80
all shell
shell=0.8mm
shell=1mm
shell=1.2mm
60
shell=1.4mm
shell=1.6mm
40
20
0
0
20
40
60
80
100
Shell MOR (MPa)
Composite MOR (MPa)
100
all core
(c)
80
all shell
shell=0.8mm
shell=1mm
shell=1.2mm
60
shell=1.4mm
shell=1.6mm
40
20
0
0
20
40
60
80
100
Shell MOR (MPa)
Figure 4.8 Comparison of MOR for coextruded composites with all core- and all shellcomposites: (a) weak core, (b) moderate core, and (c) strong core.
82
Composite Impact (kJ/m2)
35
All Core
All Shell
shell=0.8mm
shell=1mm
shell=1.2mm
shell=1.4mm
shell=1.6mm
(a)
30
25
20
15
10
5
0
0
5
10
15
20
25
30
35
Shell Impact (kJ/m2)
Composite Impact(kJ/m2)
35
All Core
All Shell
shell=0.8mm
shell=1mm
shell=1.2mm
shell=1.4mm
shell=1.6mm
(b)
30
25
20
15
10
5
0
0
5
10
15
20
25
30
35
Shell Impact (kJ/m2)
Composite Impact (kJ/m2)
35
All Core
All Shell
shell=0.8mm
shell=1mm
shell=1.2mm
shell=1.4mm
shell=1.6mm
(c)
30
25
20
15
10
5
0
0
5
10
15
20
25
30
35
Shell Impact (kJ/m2)
Figure 4.9 Comparison of impact strengths for coextruded composites with all core- and all
shell-composites: (a) weak core, (b) moderate core, and (c) strong core.
83
Table 4.3 Two-way ANOVA tests of the effect of GF content in shell and shell thickness on
mechanical properties of different core coextruded composites.
Core
Type
Weak
Moderate
Strong
Variables
GF
Thickness
GF × Thickness
GF
Thickness
GF × Thickness
GF
Thickness
GF × Thickness
Property
MOEa (GPa)
MOR (MPa)
IM (kJ/m2)
< 0.0001
< 0.0001
< 0.0001
< 0.0001
0.0447
< 0.0001
< 0.0001
< 0.0001
< 0.0001
< 0.0001
< 0.0001
0.0027
< 0.0001
< 0.0001
< 0.0001
< 0.0001
0.0053
< 0.0001
< 0.0001
< 0.0001
< 0.0001
< 0.0001
< 0.0001
< 0.0001
0.2531
< 0.0001
0.0002
a
The values shown are p-values of 2-way ANOVA tests. The p-values smaller than 0.05 indicate
significant influences of the corresponding treatments on each property at the 95% confidence
level.
4.4 CONCLUSIONS
The effects of various GF contents in a shell layer and shell thickness changes on the
flexural and impact properties of coextruded WPCs in combination with three core systems (low,
moderate, and strong) were studied in this work.
GF fillers in the shell layer behaved as an effective reinforcement for the whole corecoextruded composites with the increase of GF content. The comparisons of flexural property of
coextruded composites with different core qualities show that GF reinforcements were optimized
at high GF loading levels. Also, GF alignments in the shell layer of the coextruded systems as
shown in SEM micrographs played an important role in determining the flexural property. Since
GFs were well bonded to the plastic matrix in the shell layer, perpendicularly-aligned GFs to the
applied flexural stress effectively improved the flexural property of coextruded composites with
GF-filled shell. However, in impact testing, the GF alignment was ineffectively activated and led
to no strength improvements.
84
For the effect of shell thickness, the flexural property of the whole composites increased
with the increase of shell MOE and MOR at a given shell thickness. When the flexural property
of shell was less than that of core, the increase of shell thickness led to reduced flexural property.
Thus, thinner shell should be preferred for this system. On the other hand, when the flexural
property of shell was higher than that of core, the opposite was true. Thus, thick shell should be
preferred for this system.
4.5 REFERENCES
Charles, A.H. 2000. Modern Plastics Handbook: McGraw-Hill.
DiBenedetto, A.T. 2001. Tailing of interfaces in glass fiber reinforced polymer composites: a
review. Materials Science and Engineering: A 302: 74-82.
Fu, S.-Y., and B. Lauke. 1996. Effect of fiber length and fiber orientation distributions on the
tensile strength of short-fiber-reinforced polymers. Composites Science and Technology 56:
1179-90.
Fu, S.Y., B. Lauke, E. Mader, X. Hu, and C.Y. Yue. 1999. Fracture resistance of short-glassfiber-reinforced and short-carbon-fiber-reinforced polypropylene under Charpy impact load and
its dependence on processing. Journal of Materials Processing Technology 90: 501-7.
Ghaus, M.R., and S. Hamid. 2008. Glass-fiber-reinforced wood/plastic composties. Journal of
Vinyl and Additive Technology 14: 39-42.
Gupta, V.B., R.K. Mittal, P.K. Sharma, G. Mennig, and J. Wolters. 1989. Some Studies on Glass
Fiber-Reinforced Polypropylene .1. Reduction in Fiber Length during Processing. Polymer
Composites 10: 8-15.
Hine, P.J., S.W. Tsui, P.D. Coates, I.M. Ward, and R.A. Duckett. 1997. Measuring the
development of fibre orientation during the melt extrusion of short glass fibre reinforced
polypropylene. Composites Part a-Applied Science and Manufacturing 28: 949-58.
Jiang, H.H., D.P. Kamdem, B. Bezubic, and P. Ruede. 2003. Mechanical properties of poly(vinyl
chloride)/wood flour/glass fiber hybrid composites. Journal of Vinyl & Additive Technology 9:
138-45.
Jiang, L., M.P. Wolcott, J.W. Zhang, and K. Englund. 2007. Flexural properties of surface
reinforced wood/plastic deck board. Polymer Engineering and Science 47: 281-8.
Jin, S., and L.M. Matuana. 2010. Wood/plastic composites co-extruded with multi-walled carbon
nanotube-filled rigid poly(vinyl chloride) cap layer. Polymer International 59: 648-57.
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Kacir, L., M. Narkis, and O. Ishai. 1975. Oriented Short Glass-Fiber Composites .1. Preparation
and Statistical-Analysis of Aligned Fiber Mats. Polymer Engineering and Science 15: 525-31.
Kim, B.-J., and Q.L. Wu. 2012. Mechanical and Physical Properties of Core-Shell Structured
Wood Plastic Composites: Effect of Shells with Hybrid Mineral and Wood Fillers. Composite
Part B, In review.
Nunez, A.J., P.C. Sturm, J.M. Kenny, M.I. Aranguren, N.E. Marcovich, and M.M. Reboredo.
2003. Mechanical characterization of polypropylene-wood flour composites. Journal of Applied
Polymer Science 88: 1420-8.
Rothon, R.N. 1999. Mineral fillers in thermoplastics: Filler manufacture and characterisation.
Mineral Fillers in Thermoplastics I 139: 67-107.
Shakeri, A., and M. Raghimi. 2010. Studies on Mechanical Performance and Water Absorption
of Recycled Newspaper/Glass Fiber-reinforced Polypropylene Hybrid Composites. Journal of
Reinforced Plastics and Composites 29: 994-1005.
Stark, N.M., and L.M. Matuana. 2007. Coating WPCs using co-extrusion to improve durability.
In Conference for Coating Wood and Wood Composites: Designing for Durability, 1-12. Seattle,
WA,.
Thomason, J.L. 2002. The influence of fibre length and concentration on the properties of glass
fibre reinforced polypropylene: 5. Injection moulded long and short fibre PP. Composites Part aApplied Science and Manufacturing 33: 1641-52.
Thwe, M.M., and K. Liao. 2002. Effects of environmental aging on the mechanical properties of
bamboo-glass fiber reinforced polymer matrix hybrid composites. Composites Part a-Applied
Science and Manufacturing 33: 43-52.
Truckenmuller, F., and H.G. Fritz. 1991. Injection-Molding of Long Fiber-Reinforced
Thermoplastics - a Comparison of Extruded and Pultruded Materials with Direct Addition of
Roving Strands. Polymer Engineering and Science 31: 1316-29.
Tselios, C., D. Bikiaris, P. Savidis, C. Panaviotou, and A. Larena. 1999. Glass-fiber
reinforcement of in situ compatibilized propylene/polyethylene blends. Journal of Materials
Science 34: 385-94.
Tungjitpornkull, S., K. Chaochanchaikul, and N. Sombatsompop. 2007. Mechanical
characterization of E-chopped strand glass fiber reinforced Wood/PVC composites. Journal of
Thermoplastic Composite Materials 20: 535-50.
Velmurugan, R., and V. Manikandan. 2007. Mechanical properties of palmyra/glass fiber hybrid
composites. Composites Part a-Applied Science and Manufacturing 38: 2216-26.
Wypych, G. 2010. Handbook of fillers, 3th ed: ChemTec Publishing.
Xanthos, M. 2005. Functional fillers for plastics: Wiley-Vch Co.
86
Yao, F., and Q.L. Wu. 2010. Coextruded Polyethylene and Wood-Flour Composite: Effect of
Shell Thickness, Wood Loading, and Core Quality. Journal of Applied Polymer Science 118:
3594-601.
87
CHAPTER 5 SOUND TRANSMISSION PROPERTIES OF MINERAL FILLED HIGH
DENSITY POLYETHYLENE (HDPE) AND WOOD-HDPE COMPOSITES
5.1 INTRODUCTION
Noise is an unwanted sound which is accompanied by industrial machinery and high
speed transportations (Jayaraman 2005; Yilmaz 2009). For example, noise level near a highway
amounts to the 70 dB level, and the level is increased up to the 80 dB level at heavily used urban
road intersections (Kotzen and English 1999). Excessive noises frequently affect people by
fatigue, reduced working efficiency and health system disorders (Suzanne et al. 2006). Various
barrier materials including concrete, brick, metal, plastic, wood and composites have been used
to reduce the noise levels. Among the materials, viscoelastic polymer materials show great
potential for dampening sound and vibration. However, most polymer materials have lower
elastic modulus and surface density, leading to poor sound insulating performances when solely
used as sound barriers (Wang et al. 2011).
Significant work has recently been done to investigate acoustic properties of filled plastic
composites. Among the studies, Lee et al. (Lee et al. 2010) investigated the tensile and sound
transmission loss (TL) properties of composites made of acrylonitrile butadiene styrene (ABS)
and carbon-black. It was shown that 5% carbon-black filled composite had more than 10% TL
increase compared with unfilled controls. Also, the composites with a high elastic modulus
showed higher TL than the materials with a low elastic modulus. It was suggested that increasing
tensile modulus through reinforcing the composite structure could be an effective technique for
blocking the sound energy. Lee et al. (Lee et al. 2008) studied the sound insulation effects of
carbon-nanotube (CNT) filled ABS composites. As CNT content in ABS composites increased,
the TL values were improved. At the frequency level of 3400 Hz, the average TL values of
88
composites with 15 vol.% of CNT were 21.7% higher than that of ABS-only composite. It was
also found that the increased stiffness from the use of CNTs played more important role on the
TL improvement of CNT-filled composites compared with original TL mass law. Costana et al.
(Cotana et al. 2007) evaluated TL properties of concrete composites filled with different
densities of pumice, lapillus and rubber. In here, composite compositions made a noticeable
influence on the TL properties. e.g., TL values of lapillus/pumice mixture filled composites were
higher than those of lapillus-only and pumice-only filled composites. Among mixtures, the
lapillus/pumice filled composite showed higher TL value compared with the pumice/rubber
filled composite. This was due to the density changes of composites with different densities of
fillers. Wang et al. (Wang et al. 2011) studied TL of laminated mica-filled poly (vinyl chloride)
composites. In their study, increased TL and resonance frequency were reported with the
increase of mica content. They suggested that stiffness and surface density were important
factors influencing on the TL properties of mica filled composites and the sound insulation
properties of the composites were well described by stiffness and mass law.
Wood plastic composites (WPCs), as new generation green composites, offer advantages
in relatively light weight, excellent recyclability, low toxicity and high thermal stability. Thus,
their application as noise barrier uses can offer a competitive alternative to the conventional
noise barriers. For this purpose, acoustic properties of WPCs, influenced by composite
formulations need to be fully understood.
The objectives of this study were firstly to develop an experimental procedure for
studying sound transmission properties of solid plastic and wood plastic composites using an
impedance tube method and secondly to investigate the effect of mineral type and loading levels
on TL properties of the composites in comparison with various corresponding TL predictions.
89
5.2 THEORETICAL APPROACH
5.2.1 Impedance Tube
The impedance tube system for sound transmission measurements is schematically
illustrated in Figure 5.1 (Top). TL measurement is done through four microphones positioned at
up and down streams of a test sample in Figure 5.1 (Bottom). In this system, sound pressures at
the four measurement locations x1 to x4 can be expressed as super-positions of positive and
negative directed plane waves (±jkx) (Olivieri, Bolton, and Yoo 2007; Song and Bolton 2000):
P1
Ae
jkx1
Be jkx1
P2
Ae
jkx2
Be jkx2
P3
Ce
jkx3
De jkx3
P4
Ce
jkx4
De jkx4
(5.1)
Where, k is the wave number in ambient air and A to D are complex plane wave
coefficients, representing the noise amplitudes. This equation can be rearranged to solve the
respective coefficients in terms of the four sound pressures (P1 to P4) as:
A
j ( P1e jkx2 P2 e jkx1 )
2 sin k ( x1 x2 )
B
j ( P2 e jkx1 P1e jkx2 )
2 sin k ( x1 x2 )
C
j ( P3e jkx4 P4 e jkx3 )
2 sin k ( x3 x4 )
D
j ( P4 e jkx3 P3e jkx4 )
2 sin k ( x3 x4 )
(5.2)
Therefore, by measuring the sound pressures at locations x1 to x4, the coefficients (A to D)
are determined and these coefficients provide the input data for subsequent transfer matrix
calculations (Olivieri, Bolton, and Yoo 2007; Song and Bolton 2000; Yousefzadeh et al. 2008).
90
Figure 5.1 Schematic diagram of hardware setup (top) and the TL measurement in an
impedance tube (bottom).
91
The transfer matrix, relating with sound pressures (P) and particle velocities (V) at the
two surfaces (front and rear) of the test sample, extending from x=0 (front) to x=d (rear), has the
following form (Olivieri, Bolton, and Yoo 2006, 2007):
T11
P
V
T12
P
T21 T22 V
x 0
(5.3)
x d
Where, Tij are frequency-dependent quantities, related to the acoustical properties of the
test sample. Thus, the P and V at the two surfaces of the test sample can be effectively expressed
by the positive and negative plane wave components (± jkx and complex coefficients):
Px
V
x 0
Px
V
0
d
A B
(5.4a)
A B
0c
(5.4b)
Ce
jkd
De jkd
(5.4c)
Ce
jkd
De jkd
0c
(5.4d)
x d
Where, ρ0 is ambient air density and c is sound speed in air. When the plane wave
components are known, based on measurements of the complex pressures at the four locations,
the P and V values at the two surfaces of the test sample can be determined. However, Eq. (5.3)
represents two equations for four unknowns (i.e., T11, T12, T21 and T22). Thus, in order to solve
the transfer matrix elements, two additional equations are required. The additional equations are
generated by making one more measurement after changing the termination of the impedance
tube from open to close. In a matrix form, the result of the two independent measurements (open
and close) is expressed as (Song and Bolton 2000):
PO
PC
VO VC
T11
x 0
T12
PO
PC
T21 T22 VO VC
(5.5)
x d
92
Where, the subscripts, O (open) and C (close) on P and V denote the two different
termination conditions. This approach is referred to as a two-load method (Munjal 1987; Olivieri,
Bolton, and Yoo 2006; Song and Bolton 2000) and the above transfer matrix elements are easily
calculated by inverting the latter expression to obtain
T11
T12
1
T21 T22
PO
VO
PO
x
V
0 C
x
V
0 C
x d
x d
PC
x d
VC
VC
x d
x
V
0 O
x
V
0 O
PC
VO
x d
x d
PO
x d
PO
x d
x
P
0 C
x
V
d C
x d
PC
x 0
PC
(5.6)
x
P
0 O
x d
x
V
d O
x 0
Pierce (Pierce 1981) suggested that for symmetrical systems, T11 = T22. Olivieri et al.
(Olivieri, Bolton, and Yoo 2006) further reported that for symmetric homogeneous and isotropic
porous material, when T11 = T22, it is possible to get the reciprocal nature of the layer to generate
two additional equations without one more measurement:
T11
(5.7a)
T22
T11T22
T12T21
(5.7b)
1
By combining Eqs. (5.3), (5.7a) and (5.7b), the transfer matrix elements for a test sample
satisfying the above conditions can be expressed directly by the sound pressures (P) and particle
velocities (V) at the two surfaces of the test sample for one termination condition:
T11
T12
1
T21 T22
P x dV
V
P x 0V
x d
2
x 0
x d
P x 0V
V
2
x d
P x dV
x 0
2
x 0
Px
P x dV
(5.8)
2
0
x d
Px
d
P x 0V
x 0
This approach is called as a one-load method, differed from the two-load method using
two measurements (Song and Bolton 2000). Once the transfer matrix elements of the test sample
are obtained by using either the one- or two-load methods, all of the other acoustical properties
can be acquired.
93
When a depth (d) of a test sample backed by an anechoic termination is considered, it can
be assumed that D=0, the P and V at the two surfaces of the test sample can be expressed as:
Px
V
x 0
Px
V
0
d
1 Ra
(5.9a)
1 Ra
0c
(5.9b)
jkd
Ta e
Ta e
x d
0
(5.9c)
jkd
(5.9d)
c
Where, reflection coefficient (Ra) = B/A and transmission coefficient (Ta) = C/A for an
anechoic termination. After substituting Eqs. (5.9a), (5.9b), (5.9c) and (5.9d) into Eq. (5.3), Ra
and Ta can be expressed as, respectively:
T11
Ra
T11
T12
0c
T12
0c
0
cT21 T22
(5.10a)
0
cT21 T22
2e jkd
Ta
T11
(5.10b)
T12
T21 T22
0c
The normal incident TL of a sample can be calculated from Eq. (5.10b) as:
TLnormal ( dB ) 10 log 10
1
Ta
(5.11)
2
Where, the subscript, normal means that the sample is exposed to a normally incident
sound wave field. When the same ambient air is on the two surfaces of the sample, Ta
2
is the
normal incident power transmission coefficient for the sample with an anechoic termination.
This TL is typically determined according to standard procedures (ASTM 2002), where the
sample is exposed to a diffuse sound field. In this study using a perfectly anechoic termination
94
(i.e., D=0) as mentioned above, the TL can be also expressed as (Olivieri, Bolton, and Yoo 2007;
Wang, He, and Geng 2008):
TLnormal (dB) 10 log 10
A( anechoic)
C ( anechoic)
2
(5.12)
When Eqs. (5.9a), (5.9b), (5.9c) and (5.9d) with a perfectly anechoic termination (i.e.,
D=0) are substituted into Eq. (5.3), the TL is calculated to be:
TLnormal (dB) 10 log 10
1
T11
4
T12
0c
2
0
cT21
T22
(5.13)
Eq. (5.13) is incorporated in the B&K acoustic testing program to compute the TL values
of test samples. In general, the TL is more easily expressed in a logarithmic form as:
TLnormal (dB) 10 log 10
Ii
It
(5.14)
Where, Ii is the incident sound intensity on the surface of the sample, and It is the
transmitted sound intensity through the sample.
5.2.2 Mass Law
When incident sound waves hit a sample, the test sample vibrates due to the changes of
ambient sound pressures. This vibration energy dissipates in and out of the sample during sound
transmission process and the dissipation increases with the increase of the sample mass. This
relation is called mass law (Lee et al. 2008). In the mass law, the TL value of a sound barrier
sample is calculated as (Jones 1979):
TL( , ) 10 log 10 1
m cos
2 c
2
(5.15)
95
Where, ω, m, θ and ρc are sound angular frequency, surface area density of sample per
unit area, angle of incident sound wave and characteristic impedance of medium, respectively.
The TL value calculated from the above equation is mainly controlled by ω, m, and θ with a
determined ρc value. However, a small scale of experiment using an impedance tube adopts
normal sound waves and the angle of the incident sound waves is perpendicular to the test
sample (θ=0). Thus, the above TL equation can also be expressed as:
TL(dB) normal 10 log 10 1
In Eq. (5.16),
m
2 c
2
(5.16)
2 f and this can also be expressed as the following equation for the
normal incident sound frequency (Folds and Loggins 1977; Lee and Xu 2009):
TLnormal
10 log 10 1
fm
c
2
(5.17)
Where, f represents sound frequency. According to previous researches, the 6 dB TL
increase per octave by doubling sample mass is broadly accepted.
5.2.3 Stiffness Law
The typical TL graph is composed of stiffness-controlled region (S-region) in low
frequency ranges and mass law-controlled region (M-region) in high frequency ranges with
increasing sound frequency (Irwin and Graf 1979). Hence, the stiffer a sound sample is, the
higher TL values it shows in the low frequency ranges. According to the definition (Ng and Hui
2008), the stiffness of material can be expressed as:
S
1
12
Eh 3
1 2
(5.18)
96
Where S, E, h, and ν are stiffness, modulus of elasticity (MOE), sample thickness, and
Poisson’s ratio of sample, respectively. From Eq. (5.18), it can be noticed that MOE and sample
thickness significantly affect the stiffness of materials. The MOE and sample thickness also can
lead to the changes of the speed of longitudinal sound wave (Cp) in materials and subsequent
sound resonance frequency (Rf). Cp and Rf are expressed as Eq. (5.19) and Eq. (5.20),
respectively (Wang et al. 2011):
Cp
R f ab
E
m(1
2
(5.19)
)
0.45C p h
2
a
2
b
r
(5.20)
r
Where, m in Eq. (5.19) is area density of sample and r, a and b in Eq. (5.20) are radius of
sample surface, and integers (1, 2, …) from the order of sound wave resonance, respectively.
Berry and Nicolas (Berry and Nicolas 1994) represented the TL with stiffness-controlled region
by using Eq. (5.17) and Eq. (5.20) as:
TLstiffness 10 log 10
fm
c
2
1
Rf
2
2
(5.21)
f
Where Rf, f, m and ρc represent sound resonance frequency, incident sound frequency,
area density of sample per unit area and characteristic impedance of medium, respectively.
5.3 EXPERIMENTAL
5.3.1 Materials and Composite Sample Preparation
Materials and formulations of each composite sample for acoustic testing are listed in
Table 5.1.
97
Table 5.1 Materials and formulations of each composite sample for acoustic testing.
Explanation
Materials
Formulation and Content
HDPE (HGB 0760)
Pine wood flour (100 mesh)
Precipitated calcium carbonate (PCC)
Nanoclay (nanoMax®)
PE Composite 1) PCC: 0, 20, 40% + HDPE
2) Nanoclay: 0, 4, 8% + HDPE
WPC
1) PCC: 0, 20, 40% + wood: 10% + HDPE
2) Nanoclay: 0, 4, 8% + wood: 10%+ HDPE
Supplier
ExxonMobil Co., TX
American Wood Fibers, WI
Domino SugarCompany, LA
Nanocor Company, TX
Lab blend
Lab blend
Lab WPC blend
Lab WPC blend
Wood flour (WF) was dried at 85°C for 24 hr to reach moisture content level less than
2%. PCC (precipitated calcium carbonate) and clay were also dried at 85°C for 24 hr to prevent
from the effects of moisture. PCC used in this experiment had the average sizes of 1.2 µm and
each clay level blended with base HDPE matrix was based on the active clay amount, which is
included in the nanoMax-HDPE masterbatch. In the first step, a CW Brabender Intelli-torque
twin-screw extrusion machine (CW Brabender Instruments, South Hackensack, NJ) was used to
make HDPE and PCC or clay base blends (HDPE : PCC = 60 : 40; HDPE : clay = 92 : 8). After
that, the pellets were diluted with HDPE or blended with WF to meet the pre-determined mixing
ratios and then extruded with a Leistritz Micro-27 co-rotating parallel twin-screw extruder
(Leistritz Corporation, Allendale, NJ) in the second step. Materials and formulations for sample
manufacturing are listed in Table 1. The blending temperature profiles were 155, 165, 175, 180,
180, 180, 180, 175, 170, 160, and 175°C for PCC filled composites and 155, 165, 170, 175, 175,
175, 175, 170, 160, 155, and 170°C for clay filled composites from feeding zone to die. The
extruder rotation speed was 60 rpm. The extruded blends were pelletized again and then dried in
an oven at 85°C for 24 hr. The prepared pellets were compression molded by using Wabash
V200 hot press (Wabash, ID, USA).
98
Thirty ton compression force was applied to composite plates at 180oC for 10 min and
then the plates were cooled down to room temperature under the same pressure. The
manufactured composite plates were water-jetted with a milling machine to produce acoustic
samples with target size of 29 (diameter) x 9 (thickness) mm.
5.3.2 Mechanical Property Characterization
Three-point flexural test was performed using a model 5582 Instron testing machine
(Instron Inc., Norwood, MA) at a crosshead speed of 3 mm/min. Each sample with a size of 125
(length) ×15 (width) ×9 (thickness) mm was prepared from the same compression-molded
composite plates as the acoustic samples. Tinius Olsen Mode 1892 impact tester (Tinius Olsen
Inc., Horsham, PA) was used to test un-notched Izod impact strength according to the ASTM
D256. The samples having 3-mm thickness were prepared for the impact strength. These samples
were also acquired from the same compression-molded composite panels as the acoustic samples
by cross-cutting.
5.3.3 Acoustic Property Characterization
To measure TL values of prepared acoustic samples, Brüel & Kjær impedance tube
instrument composed of power amplifier (Type 2716C), transmission loss tube kit (Type 4206T),
signal amplifier (Type 3560C-S29) and pulse FFT analysis program (Labshop ver. 16.01) was
used as shown in Figure 5.1(Top). Small tube (500-6400 Hz) set-ups were adopted. Prior to
acoustic testing, test samples were sanded around the perimeter such that they fit into the test
tube without much lateral force from the tube wall and also, petroleum jelly (Vaseline) was
applied to the sample edges to prevent from sound pressure leaks. Final sample dimension and
weight were measured and used to calculate sample area density. To get a better TL curves
without noise fluctuations, FFT curves were smoothed by 1/3 octave band width.
99
5.4 RESULTS AND DISCUSSION
5.4.1 Basic Properties of Mineral Filled HDPE and WPC
Densities of clay, PCC and HDPE are, respectively, 1.60, 2.60, and 0.96 g/cm3 (Kim et al.
2012; Klyosov 2007; Wypych 2010). Thus, the addition of clay and PCC in the HDPE matrix
increased the surface area density of the composites for the same sample size (Table 5.2). Pure
HDPE had an area density of 8.59 kg/m2. The density increased to 8.95 kg/m2 at the 8% clay
loading level and to 10.83 kg/m2 at the 40% PCC level. Wood plastic composites (10% WF)
filled with clay and PCC had much similar area densities compared to corresponding HDPE
composite with clay and PCC only, respectively. Thus, the use of 10% WF did not alter the
composite area density much.
Table 5.2 Mechanical properties (MOE, MOR, and impact strength), stiffness, and area
density of mineral filled composites and WPCs as a function of filler content.
System
Clay Only
PCC Only
Clay + WF
PCC + WF
a
Filler (%)
MOEa
(GPa)
MOR
(MPa)
0
4
8
20
40
0 + 10
4 + 10
8 + 10
20 + 10
40 + 10
1.20 (0.02)
1.31 (0.01)
1.34 (0.05)
1.47 (0.05)
1.51 (0.04)
1.39 (0.03)
1.45 (0.06)
1.47 (0.06)
1.64 (0.05)
1.49 (0.04)
41.57 (1.21)
33.17 (0.05)
30.42 (0.98)
46.20 (0.56)
32.49 (0.47)
37.47 (0.82)
37.67 (1.04)
33.22 (0.39)
37.14 (1.50)
34.33 (0.17)
Impact
Strength
(kJ/m2)
No Break
5.90 (0.11)
5.18 (0.19)
11.03 (2.35)
7.00 (0.65)
17.98 (0.89)
10.19 (2.21)
6.58 (1.50)
13.50 (2.74)
4.37 (0.36)
Numbers in parenthesis are standard deviations.
100
Stiffness
of Unit Area
(×105N/m2)
119.05
129.47
131.94
144.47
148.41
136.61
141.99
143.43
159.73
145.12
Area Density
of Unit Area
(kg/m2)
8.59
8.79
8.95
9.75
10.83
8.26
8.47
8.68
9.57
10.66
The neat HDPE samples had an MOE value of 1.2 GPa. At the 4 and 8% clay loading
levels, the modulus increased to 1.31 GPa and 1.34 GPa, respectively. Early work in this field
(Lei et al. 2007) attributed the modulus increase to clay exfoliation to form nano-platelets in the
plastic matrix. The extent of exfoliation depends on mixing and the use of coupling agents. PCC
filled HDPE also showed higher modulus compared with neat-HDPE, probably due to higher
filler content (Kim et al. 2012). The corresponding stiffness for both types of composites also
increased with increased clay and PCC loading levels. On the other hand, both bending MOR
and impact strength decreased with increased clay and PCC loading levels. Possible particle
aggregation in the HDPE matrix, which creates stress concentration at particle-matrix interface
led to decreased overall composite strength.
For clay filled WPC, increased clay loading led to some improvement in the bending
MOE (e.g., an increment from 1.39 GPa for 0% clay level to 1.47 GPa for 8% clay level), similar
to clay-filled HDPE composites. There was an overall modulus value enhancement compared
with corresponding HDPE-only composites, showing some synergetic effect of clay and wood
fibers. For PCC-filled WPC, MOE was obviously increased at the 20% PCC level compared with
PCC-HDPE composite, but at the 40% PCC level, the MOE was somewhat reduced. There was a
decreasing trend for strength properties (especially, impact strength) with the increased clay or
PCC loading level. The reduced strength properties of filled composite are attributed to the
aggregation effect of filler particles in the HDPE matrix (Kim et al. 2012).
5.4.2 General TL Curves of Filled HDPE and WPCs
Figure 5.2 shows a comparison of measured TL-frequency curves for clay (a) and PCC (b)
filled HDPE and WPC material.
101
Transmission Loss (dB)
70
(a)
65
PE
PE/Clay4
PE/Clay8
WPC
WPC/Clay4
WPC/Clay8
60
55
50
45
Low Frequency
High Frequency
40
0
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
Transmission Loss (dB)
70
(b)
65
PE
PE/PCC20
PE/PCC40
WPC
WPC/PCC20
WPC/PCC40
60
55
50
45
High Frequency
Low Frequency
40
0
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
Figure 5.2 Comparisons of TL curves between clay-filled HDPE and WPCs (a) and
between PCC filled HDPE and WPCs (b).
102
The TL curves shifted upwards as clay or PCC level increased, indicating improved TL
properties at higher filler loading levels. The area density of clay filled HDPE was only slightly
higher than that of neat-HDPE (Table 5.2). The observed TL improvement at higher clay loading
levels, especially in a low frequency range, probably indicates the effect of nano-clay in
composites. The clay particles, after intercalation and exfoliation, had platelet shapes with
thickness in the nanometer range, and they were distributed in the plastic matrix (Ataeefard and
Moradian 2011; Gopakumar et al. 2002; Ton-That et al. 2004). The platelets can help deflect
sound wave and increase its pathway with energy dissipations as heat, leading to increased TL
value (Bharadwaj 2001; Lee et al. 2008; Picard et al. 2007). For PCC filled composites, the
addition of PCC in HDPE led to distinctively increased area density and the density increases
consequently improved TL values of PCC-filled composites in comparison with neat-HDPE.
However, the 40% PCC filled composite showed a lower TL curve in the low frequency range
and higher TL curve in the high frequency range, compared with the 20% PCC filled composite
as shown in Figure 5.2(b).
The incorporation of 10% WF in the clay and PCC filled composites led to the overall
improvements of the composite TL values in comparison with corresponding HDPE composites.
The results show a synergistic effect of clay or PCC and wood particles in the composites
through blocking/reflecting the sound wave and enhancing composite stiffness. The initial large
difference of the observed TL values in the low and middle frequency ranges indicates stiffness
controls in the TL of the composites. For PCC-filled composites, composite stiffness was lower
at the 40% loading level in comparison with 20% PCC loading. Measured TL values also
decreased with PCC loading increases from 20% to 40%.
103
Transmission Loss (dB)
80
HDPE (Experimental)
HDPE (Mass Law)
HDPE (Stiffness-1)
HDPE (Stiffness-2)
(a)
70
60
50
40
30
20
10
0
0
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
80
(b)
80
PE/Clay4 (Experimental)
PE/Clay4 (Mass Law)
PE/Clay4 (Stiffness-1)
PE/Clay4 (Stiffness-2)
70
Transmission Loss (dB)
Transmission Loss (dB)
90
60
50
40
30
20
10
0
(c)
70
PE/Clay8 (Experimental)
PE/Clay8 (Mass Law)
PE/Clay8 (Stiffness-1)
PE/Clay8 (Stiffness-2)
60
50
40
30
20
10
0
0
1000
2000
3000
4000
5000
6000
7000
0
1000
Frequency (Hz)
70
(d)
PE/PCC20 (Experimental)
PE/PCC20 (Mass Law)
PE/PCC20 (Stiffness-1)
PE/PCC20 (Stiffness-2)
60
3000
4000
5000
6000
7000
Frequency (Hz)
50
40
30
20
10
0
80
Transmission Loss (dB)
Transmission Loss (dB)
80
2000
PE/PCC40 (Experimental)
PE/PCC40 (Mass Law)
PE/PCC40 (Stiffness-1)
PE/PCC40 (Stiffness-2)
(e)
70
60
50
40
30
20
10
0
0
1000
2000
3000
4000
5000
6000
7000
0
Frequency (Hz)
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
Figure 5.3 Comparisons of experimental TL curves of filled HDPE in comparison with
mass and stiffness law TL predictions: (a) HDPE, (b) HDPE/Clay4, (c) HDPE/Clay8, (d)
HDPE/PCC20, and (e) HDPE/PCC40.
104
80
WPC (Experimental)
WPC (Mass Law)
WPC (Stiffness-1)
WPC (Stiffness-2)
Transmission Loss (dB)
(a)
70
60
50
40
30
20
10
0
0
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
70
WPC/Clay4 (Experimental)
WPC/Clay4 (Mass Law)
WPC/Clay4 (Stiffness-1)
WPC/Clay4 (Stiffness-2)
(b)
60
50
40
30
20
10
80
Transmission Loss (dB)
Transmission Loss (dB)
80
0
WPC/Clay8 (Experimental)
WPC/Clay8 (Mass Law)
WPC/Clay8 (Stiffness-1)
WPC/Clay8 (Stiffness-2)
(c)
70
60
50
40
30
20
10
0
0
1000
2000
3000
4000
5000
6000
7000
0
1000
Frequency (Hz)
70
WPC/PCC20 (Experimental)
WPC/PCC20 (Mass Law)
WPC/PCC20 (Stiffness-1)
WPC/PCC20 (Stiffness-2)
(d)
60
3000
4000
5000
6000
7000
Frequency (Hz)
50
40
30
20
10
0
80
Transmission Loss (dB)
Transmission Loss (dB)
80
2000
WPC/PCC40 (Experimental)
WPC/PCC40 (Mass Law)
WPC/PCC40 (Stiffness-1)
WPC/PCC40 (Stiffness-2)
(e)
70
60
50
40
30
20
10
0
0
1000
2000
3000
4000
5000
6000
7000
0
Frequency (Hz)
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
Figure 5.4 Comparisons of experimental TL curves of filled WPCs in comparison with
mass and stiffness law TL predictions: (a) WPC, (b) WPC/Clay4, (c) WPC/Clay8, (d)
WPC/PCC20, and (e) WPC/PCC40.
105
5.4.3 Comparison of TL with Mass and Stiffness Law Predictions
Experimental TL curves of clay and PCC filled HDPE and WPCs are shown in Figure 5.3
and Figure 5.4, respectively, in comparison with the corresponding mass law and stiffness law
TL prediction lines. The mass law TL equation used was a modified form of Eq. (5.7) (Lee and
Xu 2009):
TLnormal
20 log 10 ( fm) 42.5
(5.22)
Where, f and m represent sound frequency and area density of sample per unit area,
respectively. Stiffness TL equation was adapted from Eq. (5.21) and the first (a=1, and b=1) and
second (a=1 and b=2) orders of sound resonance frequencies were used to get two stiffness TL
prediction curves (i.e., stiffness-1 and stiffness-2).
There were initial fluctuations in the experimental TL curves at a very low frequency
range. The TL values decreased as frequency increased and then increased after passing a certain
frequency level (designated as experimental resonance frequency – RFE). The stiffness-1 and
stiffness-2 TL curves for each sample have two resonance frequencies RFS1 and RFS2,
respectively. The mass law TL curves formed linear lines after an initial rapid TL increase at the
low frequency range. The effect of filler loading and filler type on the shape of the TL curves
seems to be small. A comparison among the experimental, predicted stiffness and mass law TL
curves indicates well-defined S-region (left) and M-region (right).
The experimental TL curves of filled HDPE and WPCs (Figures 5.3(a) and 5.4(a)) show
similar trends to their corresponding mass law TL predictions for M-region. This result is similar
to the previous TL experimental data using mica-filled PVC composites with 1 mm thickness
(Wang et al. 2011). However, for S-region, there are some differences between two studies.
While the published experimental TL curves had similar features to their corresponding stiffness-
106
1 TL predictions, the experimental TL curves from the current work are positioned in between
the stiffness-1 and stiffness-2 TL predictions. The different experimental TL features in S-region
between the two studies may be related with stiffness differences. When the stiffness equation as
described in Eq. (5.18) is considered, the stiffness value is significantly affected by the thickness
of test sample (i.e., 729 times difference between 1 mm and 9 mm) and the increased sample
thickness led to remarkable stiffness value differences. Thus, the increased stiffness values from
the test samples with 9-mm thickness resulted in the remarkable TL improvements in the Sregion.
The stiffness-1 and stiffness-2 TL curves form two intersection points with the mass law
TL curve at two frequency levels, RS1-M-IP and RS2-M-IP, respectively as shown in Figure 5.5
(filled HDPE) and Figure 5.6 (filled WPCs). The experimental TL curves of filled HDPE and
WPCs are better approximate to the combined TL predictions from their corresponding stiffness1 and stiffness-2 TL curves for S-region and Mass Law TL curves for M-region. Table 5.3 shows
a comparison of the RFE, RFS1, RFS2, RS1-M-IP and RS2-M-IP values for various composites. While
the experimental RFE values of filled HDPE composites are located between the corresponding
RS1-M-IP and RS2-M-IP values, those of filled WPCs are more close to the corresponding RS2-M-IP
values or located to the right side of them. From the above result, it can be noticed that the MOE
and stiffness increases of the test samples led to the experimental TL improvements in S-region
and these consequently shifted their experimental RFE values to the high frequency range.
107
80
Transmission Loss (dB)
(a)
HDPE (Experimental)
HDPE (Stiffness-1+Mass Law)
HDPE (Stiffness-2+Mass Law)
70
60
50
40
Region-S
Region-M
2000
4000
30
0
1000
3000
5000
6000
7000
Frequency (Hz)
80
80
Transmission Loss (dB)
70
60
50
40
Region-M
Region-S
PE/Clay8 (Experimental)
PE/Clay8 (Stiffness-1+Mass Law)
PE/Clay8 (Stiffness-2+Mass Law)
(c)
Transmission Loss (dB)
PE/Clay4 (Experimental)
PE/Clay4 (Stiffness-1+Mass Law)
PE/Clay4 (Stiffness-2+Mass Law)
(b)
30
70
60
50
40
Region-S
30
0
1000
2000
3000
4000
5000
6000
7000
0
1000
Frequency (Hz)
2000
3000
4000
5000
6000
7000
Frequency (Hz)
80
80
70
60
50
40
Region-S
Region-M
30
PE/PCC40 (Experimental)
PE/PCC40 (Stiffness-1+Mass Law)
PE/PCC40 (Stiffness-2+Mass Law)
(e)
Transmission Loss (dB)
PE/PCC20 (Experimental)
PE/PCC20 (Stiffness-1+Mass Law)
PE/PCC20 (Stiffness-2+Mass Law)
(d)
Transmission Loss (dB)
Region-M
70
60
50
40
Region-S
Region-M
30
0
1000
2000
3000
4000
5000
6000
7000
0
Frequency (Hz)
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
Figure 5.5 Comparisons of experimental TL curves of filled HDPE in comparison with
combined TL predictions: (a) HDPE, (b) HDPE/Clay4, (c) HDPE/Clay8, (d) HDPE/PCC20,
and (e) HDPE/PCC40.
108
80
WPC (Experimental)
WPC (Stiffness-1+Mass Law)
WPC (Stiffness-2+Mass Law)
Transmission Loss (dB)
(a)
70
60
50
40
Region-M
Region-S
30
0
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
80
80
Transmission Loss (dB)
70
60
50
40
Region-M
Region-S
WPC/Clay8 (Experimental)
WPC/Clay8 (Stiffness-1+Mass Law)
WPC/Clay8 (Stiffness-2+Mass Law)
(c)
Transmission Loss (dB)
WPC/Clay4 (Experimental)
WPC/Clay4 (Stiffness-1+Mass Law)
WPC/Clay4 (Stiffness-2+Mass Law)
(b)
30
70
60
50
30
0
1000
2000
3000
4000
5000
6000
7000
0
1000
2000
Frequency (Hz)
80
(d)
WPC/PCC20 (Experimental)
WPC/PCC20 (Stiffness-1+Mass Law)
WPC/PCC20 (Stiffness-2+Mass Law)
70
60
50
40
Region-M
Region-S
3000
4000
5000
6000
7000
Frequency (Hz)
30
Transmission Loss (dB)
80
Transmission Loss (dB)
Region-M
Region-S
40
(e)
WPC/PCC40 (Experimental)
WPC/PCC40 (Stiffness-1+Mass Law)
WPC/PCC40 (Stiffness-2+Mass Law)
70
60
50
40
Region-S
Region-M
30
0
1000
2000
3000
4000
5000
6000
7000
0
Frequency (Hz)
1000
2000
3000
4000
5000
6000
7000
Frequency (Hz)
Figure 5.6 Comparisons of experimental TL curves of filled WPCs in comparison with
combined TL predictions: (a) WPC, (b) WPC/Clay4, (c) WPC/Clay8, (d) WPC/PCC20, and
(e) WPC/PCC40.
109
Table 5.3 Summary of sound transmission property of filled HDPE and WPCs.
Predicted Resonance
Frequency
RFS1
RFS2
(Hz)
(Hz)
Stiffness-Mass Law Curves
Intersection Points
RS1-M-IP
RS2-M-IP
(Hz)
(Hz)
Sample
Name
Experimental
Resonance
Frequency
RFE (Hz)
HDPE
HDPE/Clay4
HDPE/Clay8
HDPE/PCC20
HDPE/PCC40
3216
3592
3792
3564
3416
2040
2168
2169
2179
2016
5099
5421
5422
5448
5041
1448
1540
1540
1548
1432
3624
3852
3852
3872
3580
WPC
WPC/Clay4
WPC/Clay8
WPC/PCC20
WPC/PCC40
4080
4572
4980
4588
3524
2435
2467
2432
2455
2002
6087
6169
6079
6138
5006
1728
1752
1728
1744
1424
4324
4380
4320
4360
3556
5.4.4 Effect of Filler Loading Level on Mean TL
For a more comprehensive comparison of the sound insulation ability of clay- and PCCfilled HDPE and WPCs with various filler contents, their averaged TL (A-TL) values in both Sregion and M-region are plotted in Figure 5.7. The A-TL values of clay filled HDPE and WPCs
increased with increased clay content in both S-region and M-region (Figure 5.7(a)). While the
clay filled HDPE had more obviously improved A-TL values only at the 8% clay level, the clay
filled WPCs showed enhanced A-TL values even at the 4% clay level, indicating synergistic
effect of WF.
The PCC-filled HDPE and WPCs showed their highest A-TL values at the 20% PCC
level. At the 20% PCC loading level, the A-TL value in S-region was much improved with the
use of 10% WF. This indicates that the stiffness increase in the 20% PCC-filled WPC played a
significant role in the observed TL values.
110
Transmission Loss (dB)
60
(a)
TL-HDPE (S)
TL-WPC (S)
TL-HDPE (M)
TL-WPC (M)
55
50
45
0
4
8
Clay Content (%)
Transmission Loss (dB)
60
(b)
TL-HDPE (S)
TL-WPC (S)
TL-HDPE (M)
TL-WPC (M)
55
50
45
0
20
40
PCC Content (%)
Figure 5.7 Comparisons of Averaged TL values between clay filled HDPE and WPCs (a)
and PCC filled HDPE and WPCs (b) in S-region and M-region.
111
However, with the further increase of PCC contents to the 40% level, the uniformity of
composites becomes worse (due to the PCC particle aggregations) and the composite stiffness
was decreased, leading to decreased A-TL values. In the M-region, the A-TL values were
increased with increased PCC content due to the area density increases of composites.
5.5 CONCLUSIONS
The effect of clay and PCC on mechanical and acoustic properties of HDPE and HDPE
based WPC was studied. The experimental TL data of the composites were analyzed with the
stiffness and mass laws. The composite modulus generally increased, while the composite
strength decreased with increased filler loading level.
The stiffness and surface area density were major factors influencing the sound insulation
property of the materials. The experimental TL results show that the addition of clay or PCC
and/or WF fillers led to general RFE and TL increases of composites. However, at high filling
levels, decreased composite stiffness from filler aggregations led to reduced TL values. The
experimental TL curves of filled HDPE and WPCs are approximated with the combined TL
predictions from their corresponding stiffness-1 and stiffness-2 TL for S-region and mass law TL
for M-region.
5.6 REFERENCES
ASTM. 2002. ASTM E-90-02 Standard test method for the laboratory measurement of airborne
sound transmission loss of building partitions and elements. West Conshohocken, PA, USA:
ASTM International.
Ataeefard, M., and S. Moradian. 2011. Polypropylene/Organoclay Nanocomposites: Effects of
Clay Content on Properties. Polymer-Plastics Technology and Engineering 50: 732-9.
Berry, A., and J. Nicolas. 1994. Structural Acoustics and Vibration Behavior of Complex Panels.
Applied Acoustics 43: 185-215.
Bharadwaj, R.K. 2001. Modeling the barrier properties of polymer-layered silicate
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114
CHAPTER 6 CONCLUSIONS AND FUTURE WORK
6.1 OVERALL CONCLUSIONS
In this dissertation, the effect of mineral fillers including PCC, GF, and clay on properties
of structured wood/natural plastic composites was investigated.
The conclusions of this study are as follows:
1) The differences in chemical composition and surface morphology between RBF and GBP
led to more effective silane crosslinking on the RBF surfaces. Dry PCC particles showed
an agglomerated form with individual particles of about 1.2 micron in diamter, and
compounding PCC with R-PP/PE resin helped separate particles. Measured flexural
strength and flexural modulus of the PCC-only-filled composites increased significantly
from 15 to 30% PCC content levels, while the tensile and impact strength of composites
decreased with the addition of PCC. For composites with hybrid bamboo and PCC fillers,
tensile and flexural moduli were improved with increasing PCC content. After silane
treatment of the bamboo filler, RBF filled composites showed noticeably increased
mechanical properties compared to those of GBP filled composites. For modulus values,
PCC-bamboo-polymer composites were 3-4 times higher than those of PCC-polymer
composites at high PCC levels.
2) Significant differences existed in mechanical, WA/TS and CTE properties between
coextruded WPCs with weak and strong core systems as influenced by the shell
compositions. In the weak core system with R-PP/HDPE, coextruded composites with a
reinforced shell showed significantly improved flexural strengths compared to their coreonly composite. In the strong core system made of V-HDPE, the flexural strengths of
coextruded WPC were lowered compared with the core-only composite. Impact strengths
115
were significantly improved for both coextruded systems at low shell filling levels. High
level of filler loading in the shell led to some decreases of impact strength in both
coextruded systems. The types of impact fracture (i.e., hinge versus complete) varied
largely with core quality and filler composition in the shell. WA/TS values of coextruded
WPCs with high TPCC content in a shell layer were lower than those of core-only
composites and coextruded WPCs with high wood content in the shell layer. The use of
high percentage of plastic in a shell layer led to increased overall CTE values for
coextruded WPCs. However, increased filler loading in the shell layer led to the decrease
of CTE values for resultant coextruded WPC, especially, for the weak core system. Thus,
the combination of a relatively weak core (e.g., made of recycled plastics) and a
reinforced shell (e.g., with small sized inorganic particles) could be a cost-effective
system for coextruded WPC manufacturing.
3) GF fillers in the shell layer behaved as an effective reinforcement for the whole corecoextruded composites with the increase of GF content. The comparisons of flexural
property of coextruded composites with different core qualities show that GF
reinforcements were optimized at high GF loading levels. Also, GF alignments in the
shell layer of the coextruded systems as shown in SEM micrographs played an important
role in determining the flexural property. Since GFs were well bonded to the plastic
matrix in the shell layer, perpendicularly-aligned GFs to the applied flexural stress
effectively improved the flexural property of coextruded composites with GF-filled shell.
However, in impact testing, the GF alignment was ineffectively activated and led to no
strength increases. For the effect of shell thickness, the flexural property of the whole
composites increased with increasing shell property at a given shell thickness. When the
116
flexural property of shell was less than that of core, the increase of shell thickness led to
reduced flexural property. Thus, thinner shell should be preferred for this system. On the
other hand, when the flexural property of shell was higher than that of core, the opposite
was true. Thus, thick shell should be preferred for this system.
4) The stiffness and surface area density were major factors influencing the sound insulation
property of the materials. The experimental TL results show that the addition of clay or
PCC and/or WF fillers led to general RFE and TL increases of composites. However, at
high filling levels, decreased composite stiffness from filler aggregations led to reduced
TL values. The experimental TL curves of filled HDPE and WPCs are approximated with
the combined TL predictions from their corresponding stiffness-1 and stiffness-2 TL for
S-region and mass law TL for M-region.
6.2 FUTURE WORK
Based on limitations of this study, suggested future studies include:
(1) In PCC filled WPCs, more extensive research on various PCC particle size, shape,
and treatment are needed. These further studies could help find the effectiveness of PCC in WPC
manufacturing for not only cost reduction but also property improvement.
(2) For mineral-filled coextruded composites, more in-depth experiments are required in
a shell filling composition. These include the incorporation of different size, shape, sort, aspect
ratio, loading level, and treatment of fillers with various plastics and comparisons with various
modeling predictions based on the above filler characteristics are needed. Also, beside
conventional properties (flexural, impact, and water absorption), the evaluations for other
properties such as surface hardness, scratch resistance, fire/flame resistance, and weathering are
required for real applications.
117
(3) Sound insulation tests for mineral-filled plastic composites and WPCs with hard
surface remain at the initial stage. Hence, firstly, additional studies should be performed with
respect to the effect of fillers on the acoustic property of filled plastic composites/WPCs.
Secondly, further researches for structured composites (layered, hollow, etc) should be followed.
118
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VITA
Birm-June Kim was born in Seoul, South Korea. He attended Kookmin University, where
he received his Bachelor of Science degree in Forest Products and Biotechnology in February
1997. After that, he joined South Korean Army and worked there as a signal officer until
September 2000. He returned to research field and continued his study at the Graduate School of
Seoul National University and received his Master of Science degree in Environmental Materials
Science in February 2004. Afterwards, in August 2008, he attended Louisiana State University
for his doctoral program in the School of Renewable Natural Resources and continued his
research. Birm-June Kim will receive the degree of Doctor of Philosophy in May 2012.
125